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Acceptance Criteria of Complete Joint Penetration Steel
Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
(2019)
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ISBN 978-0-309-48044-4 | DOI 10.17226/25494
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Robert J. Connor, Curtis J. Schroeder, Bridget M. Crowley, Glenn A. Washer, and
Philip E. Fish; National Cooperative Highway Research Program; Transportation
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
N AT I O N A L C O O P E R AT I V E H I G H W AY R E S E A R C H P R O G R A M
NCHRP RESEARCH REPORT 908
Acceptance Criteria
of Complete Joint Penetration
Steel Bridge Welds
Evaluated Using Enhanced
Ultrasonic Methods
Robert J. Connor
Curtis J. Schroeder
Bridget M. Crowley
Purdue University
West Lafayette, IN
Glenn A. Washer
University of Missouri
Columbia, MO
Philip E. Fish
Fish & Associates, Inc.
Madison, WI
Subscriber Categories
Bridges and Other Structures
Research sponsored by the American Association of State Highway and Transportation Officials
in cooperation with the Federal Highway Administration
2019
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
NATIONAL COOPERATIVE HIGHWAY
RESEARCH PROGRAM
NCHRP RESEARCH REPORT 908
Systematic, well-designed, and implementable research is the most
effective way to solve many problems facing state departments of
transportation (DOTs) administrators and engineers. Often, highway
problems are of local or regional interest and can best be studied by
state DOTs individually or in cooperation with their state universities
and others. However, the accelerating growth of highway transportation results in increasingly complex problems of wide interest to highway authorities. These problems are best studied through a coordinated
program of cooperative research.
Recognizing this need, the leadership of the American Association
of State Highway and Transportation Officials (AASHTO) in 1962 initiated an objective national highway research program using modern
scientific techniques—the National Cooperative Highway Research
Program (NCHRP). NCHRP is supported on a continuing basis by
funds from participating member states of AASHTO and receives the
full cooperation and support of the Federal Highway Administration,
United States Department of Transportation.
The Transportation Research Board (TRB) of the National Academies
of Sciences, Engineering, and Medicine was requested by AASHTO to
administer the research program because of TRB’s recognized objectivity
and understanding of modern research practices. TRB is uniquely suited
for this purpose for many reasons: TRB maintains an extensive committee structure from which authorities on any highway transportation
subject may be drawn; TRB possesses avenues of communications and
cooperation with federal, state, and local governmental agencies, universities, and industry; TRB’s relationship to the National Academies is an
insurance of objectivity; and TRB maintains a full-time staff of specialists in highway transportation matters to bring the findings of research
directly to those in a position to use them.
The program is developed on the basis of research needs identified by
chief administrators and other staff of the highway and transportation
departments, by committees of AASHTO, and by the Federal Highway
Administration. Topics of the highest merit are selected by the AASHTO
Special Committee on Research and Innovation (R&I), and each year
R&I’s recommendations are proposed to the AASHTO Board of Directors and the National Academies. Research projects to address these
topics are defined by NCHRP, and qualified research agencies are
selected from submitted proposals. Administration and surveillance of
research contracts are the responsibilities of the National Academies
and TRB.
The needs for highway research are many, and NCHRP can make
significant contributions to solving highway transportation problems
of mutual concern to many responsible groups. The program, however,
is intended to complement, rather than to substitute for or duplicate,
other highway research programs.
Project 14-35
ISSN 2572-3766 (Print)
ISSN 2572-3774 (Online)
ISBN: 978-0-309-48044-4
Library of Congress Control Number 2019905465
© 2019 National Academy of Sciences. All rights reserved.
COPYRIGHT INFORMATION
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written permissions from publishers or persons who own the copyright to any previously
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Cooperative Research Programs (CRP) grants permission to reproduce material in this
publication for classroom and not-for-profit purposes. Permission is given with the
understanding that none of the material will be used to imply TRB, AASHTO, FAA, FHWA,
FMCSA, FRA, FTA, Office of the Assistant Secretary for Research and Technology, PHMSA,
or TDC endorsement of a particular product, method, or practice. It is expected that those
reproducing the material in this document for educational and not-for-profit uses will give
appropriate acknowledgment of the source of any reprinted or reproduced material. For
other uses of the material, request permission from CRP.
NOTICE
The research report was reviewed by the technical panel and accepted for publication
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The opinions and conclusions expressed or implied in this report are those of the
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Research Board; the National Academies of Sciences, Engineering, and Medicine; or the
program sponsors.
The Transportation Research Board; the National Academies of Sciences, Engineering,
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Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
The National Academy of Sciences was established in 1863 by an Act of Congress, signed by President Lincoln, as a private, nongovernmental institution to advise the nation on issues related to science and technology. Members are elected by their peers for
outstanding contributions to research. Dr. Marcia McNutt is president.
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Learn more about the National Academies of Sciences, Engineering, and Medicine at www.national-academies.org.
The Transportation Research Board is one of seven major programs of the National Academies of Sciences, Engineering, and Medicine.
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Learn more about the Transportation Research Board at www.TRB.org.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
COOPERATIVE RESEARCH PROGRAMS
CRP STAFF FOR NCHRP RESEARCH REPORT 908
Christopher J. Hedges, Director, Cooperative Research Programs
Lori L. Sundstrom, Deputy Director, Cooperative Research Programs
Waseem Dekelbab, Senior Program Officer
Megan A. Chamberlain, Senior Program Assistant
Eileen P. Delaney, Director of Publications
Natalie Barnes, Associate Director of Publications
Heidi Willis, Editor
NCHRP PROJECT 14-35 PANEL
Field of Maintenance—Area of Maintenance of Way and Structures
Alexander K. Bardow, Massachusetts DOT, Boston, MA (Chair)
Steven M. Duke, Florida DOT, Gainesville, FL
Karl H. Frank, Austin, TX
Heather E. Gilmer, Tampa Tank/Florida Structural Steel, Tampa, FL
Richard A. Pimpinella, New York State DOT, Albany, NY
Phillip W. Sauser, U.S. Army Corps of Engineers, St. Paul, MN
Hormoz Seradj, Salem, OR
Justin M. Ocel, FHWA Liaison
AUTHOR ACKNOWLEDGMENTS
The research documented in this report was performed under NCHRP Project 14-35 by the Lyles
School of Civil and Environmental Engineering at Purdue University in West Lafayette, IN. Purdue
University is the prime contractor for this study with Dr. Robert J. Connor, Professor of Civil Engineering
at Purdue, as the Project Director and Principal Investigator. The Co-Principal Investigator of this
report is Professor Glenn A. Washer from the Civil and Environmental Engineering Department at
the University of Missouri. The other authors of this report are PhD Candidate Curtis J. Schroeder and
Bridget M. Crowley, Research Assistants at the Bowen Laboratory for Large-Scale Civil Engineering
Research at Purdue University and Philip E. Fish, Senior Consultant formerly with Fish & Associates, Inc.
The authors also acknowledge the guidance and input from the Project Panel.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
FOREWORD
By Waseem Dekelbab
Staff Officer
Transportation Research Board
NCHRP Research Report 908: Acceptance Criteria of Complete Joint Penetration Steel
Bridge Welds Evaluated Using Enhanced Ultrasonic Methods presents guidelines for evaluating complete joint penetration (CJP) welds in steel bridges and proposes modifications
to the American Association of State Highway and Transportation Officials (AASHTO)/
American Welding Society (AWS) D1.5. The guidelines are based on comprehensive analytical and round robin testing programs that established the critical flaw size that would be
considered rejectable for typical bridge CJP welds and identify best practices for improved
flaw detection and flaw characterization. This report will be of immediate interest to steel
bridge fabricators and engineers.
Inspection of welds in steel bridges is necessary to ensure the quality of workmanship
during the fabrication and construction process and later on when the bridge is in service.
There are two non-destructive evaluation (NDE) methods for evaluation of complete joint
penetration (CJP) welds in steel bridges: radiographic (RT) and ultrasonic (UT). Recent
advances in enhanced ultrasonic methods, including the development of phased-array
ultrasonic technology (PAUT), allow for efficient detection and characterization of flaws
with the option of automated data collection and imaging. Criteria for categorizing weld
discontinuities as acceptable or unacceptable are codified in the AASHTO/AWS D1.5M/
D1.5: Bridge Welding Code (BWC). However, these acceptance criteria do not reflect the
full use of the capability of enhanced ultrasonic testing methods, and furthermore are not
based on the effect of weld discontinuities on bridge performance (e.g., resistance to fatigue
and fracture). In addition, some weld discontinuities that are not allowed according to
BWC are potentially not harmful and may not decrease service life. An updated acceptance
criteria based on enhanced ultrasonic testing methods for evaluation of CJP welds in steel
bridges was needed for fabricators and bridge owners.
The objectives of this research were to: (1) develop guidelines to evaluate complete joint
penetration welds in steel bridges based on updated acceptance criteria and (2) develop
proposed modifications to the BWC. The guidelines cover shop and field fabrication and
in-service evaluation and include procedures for using enhanced ultrasonic testing methods
to evaluate CJP welds in steel bridges and pertinent acceptance criteria.
Under NCHRP Project 14-35, Purdue University was asked to: (1) develop guidelines
to evaluate complete joint penetration welds in steel bridges based on updated acceptance
criteria and (2) develop proposed modifications to BWC. The guidelines address shop
and field fabrication and in-service evaluation and include procedures for using enhanced
ultrasonic testing methods to evaluate CJP welds in steel bridges and pertinent acceptance
criteria.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
A number of deliverables, provided as appendices, are not published but are available on
the TRB project website. These appendices are titled as follows:
•
•
•
•
•
•
•
•
Appendix A—T and Corner Joint Critical Flaw Size
Appendix B—Round Robin Height and Length Measurements
Appendix C—Round Robin Reported Amplitude
Appendix D—Digital RT Images of Round Robin Specimens
Appendix E—CIVA Modeling Results
Appendix F—Acoustic Property Experimental Results
Appendix G—AWS D1.5:2015 Annex K Markups
Appendix H—Directions Supplied to Round Robin Technicians
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
CONTENTS
1
Summary
4
Chapter 1 Background
4
5
5
8
8
12
14
14
24
33
33
39
44
46
53
63
65
65
69
72
72
72
74
1.1 Problem Statement and Research Objective
1.2 Scope of Study
1.3 Phased Array Ultrasonic Testing (PAUT)
Chapter 2 Research Approach
2.1 Summary of the State of the Practice
2.2 Research Methodology
Chapter 3 Findings and Application
3.1 Critical Flaw Size
3.2 Round Robin Results
3.3 CIVA Modeling
3.4 Acoustic Attenuation
3.5 Shear Wave Velocity
3.6 Calibration Blocks
3.7 Amplitude Limit for Rejection of Flaws
3.8 Amplitude Limit for Detection of Flaws
3.9 Compression Weld Acceptance Criteria
3.10 Comparison to Radiographic Testing
3.11 Technician Performance Qualification
3.12 Flaw Sizing Acceptance Criteria for Alternative UT Methods
Chapter 4 Conclusions and Suggested Research
4.1 Conclusions
4.2 Suggested Research
References
Note: Photographs, figures, and tables in this report may have been converted from color to grayscale for printing.
The electronic version of the report (posted on the web at www.trb.org) retains the color versions.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
1 SUMMARY
Acceptance Criteria of Complete Joint
Penetration Steel Bridge Welds Evaluated
Using Enhanced Ultrasonic Methods
This report summarizes the research and findings of NCHRP Project 14-35: Acceptance
Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced
Ultrasonic Methods, which focused on developing guidelines to evaluate complete joint
penetration (CJP) welds in steel bridges based on updated acceptance criteria (developed
during this research) and developing proposed modifications to the American Welding
Society (AWS) D1.5:2015 Bridge Welding Code [1]. The 2015 edition of AWS D1.5 included
Annex K, which provided an inspection procedure and acceptance criteria to apply phased
array ultrasonic testing (PAUT) to the inspection of steel bridge welds. These acceptance
criteria were workmanship-based and were carried over from previous AWS D1.1:2010
conventional ultrasonic testing (UT) methods. Rather than implement acceptance criteria
based on workmanship, this research project focused on tying the acceptance criteria
to the criticality of weld flaws using fracture mechanics in a fitness-for-service-based
approach. The 2015 edition of AWS D1.5 did not provide means for using alternative UT methods such as time-of-flight diffraction (TOFD) UT or full matrix capture
(FMC)—total focusing method (TFM) PAUT for inspection of steel bridge welds. These
methods are suited to evaluate flaw criticality based on measurements of flaw size rather
than amplitude responses.
The current version of AWS D1.5 Annex K utilizes encoded line scanning using PAUT
sector scans (S-scans) to provide sound coverage for detection and rejection of weld flaws.
Although S-scans sweep through a range of incidence angles, line scanning with a set index
offset (i.e., distance between probe and weld centerline) will only impact each point in the
volume of the weld with a single incidence angle, ignoring beam spread and reflections off
of the backwall. Therefore, the maximum amplitude determined during line scanning is
typically not maximized compared to the maximum amplitude during raster scanning. This
is a very important distinction.
Based on this fact, research was performed with the objective of developing an inspection
procedure that will diminish the scatter in results due to variability in probe location, flaw
location (i.e., transverse and through-thickness location), flaw tilt, and flaw skew. It was
determined that a two-part inspection procedure would best meet this objective. The proposed procedure would use (1) encoded line scanning to detect flaws above a set amplitude
limit and (2) inspection of these suspect locations using manual raster scanning to maximize the amplitude response for determination of acceptance/rejection. In this methodology, encoded line scanning will provide necessary sensitivity to critical flaws and provide for
permanent documentation of encoded results. Manual raster scanning is used to maximize
the amplitude response for indications which exceed the flaw detection amplitude limit to
evaluate acceptance. This testing will maximize the amplitude response of indications and
better ensure that critical flaws are rejected.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
2
An analytical parametric study was performed in order to establish the critical flaw size
that would be considered rejectable for typical bridge CJP welds, and the results were used
to identify the critical flaw size that should be reliably detected and rejected and to establish
revised acceptance and rejection criteria. The critical flaw sizes were evaluated based on
fracture mechanics using a British Standards Institute BS 7910:2013 [2] fitness-for-service
(FFS) approach. It is inherently assumed that all internal flaws are cracks. The critical size of
volumetric flaws was also evaluated through review of prior fatigue test results. This review
highlighted that the fatigue crack growth threshold of volumetric flaws is highly variable,
but when considering the lower bound, it is very similar to that of planar flaws. Therefore,
FFS analysis could be applied for volumetric flaws by evaluating them as crack-like and
determining the flaw size that will not undergo fatigue crack growth.
FFS was performed to evaluate two different failure modes for planar flaws. The first is
failure due to fatigue crack growth in order to provide infinite fatigue life. The second is
failure due to brittle fracture. The analytical studies that were performed under this task
included parametric studies of various plate geometries, welds, residual stress fields, flaw
types, flaw sizes, and locations. The results from the analytical program have identified the
target critical flaw size.
A round robin testing program was performed in order to gain insight into the capabilities
of the current technicians in the steel bridge industry and to identify best practices for
improved flaw detection and flaw characterization. The round robin testing program was
used to determine the minimum flaw size that could be reliably detected with enhanced
ultrasonic methods and how the advanced methods compare with the conventional UT
method. The round robin experimental testing program was performed by circulating weld
flaw specimens to acquire inspection data from PAUT, conventional UT, TOFD, and radiographic testing (RT) technicians. The inspection results were compared with the various
inspection methods; the comparison included the hit/miss rate, rejection rate, frequency of
false calls, accuracy of flaw height and length sizing, accuracy of flaw type characterization,
and technician variability.
The results of the round robin testing program were used to improve development of
future PAUT scanning procedures and acceptance criteria. These results also highlighted
that moving forward with the development of acceptance criteria based solely on the measurement of the flaw size—with the accuracy and reliability provided with flaw size measurements using the current PAUT workforce—is not feasible at present. Therefore, the focus
of the research was on improving on the acceptance criteria based on maximum amplitude
and flaw length in AWS D1.5 Annex K. An acceptance criteria based on physical measurements of the flaw height and length has also been provided in order to allow for alternative
UT inspection methods or future technological improvements.
Experimental testing of various bridge welds and base materials have highlighted that
the acoustic properties may vary considerably, including acoustic attenuation and shear
wave velocity. Additional calibration requirements are necessary in order to account for
possible differences between the calibration block and test object. Not properly accounting
for these acoustic properties could result in large variations of the amplitude of indications
in the test object.
The experimental results of the acoustic properties of various bridge base materials were
used to generate benchmarked material models for ultrasonic inspection simulations using
CIVA-UT. These simulations were used to develop requirements on the calibration block
material and scanning parameters in order to limit the possible error in reference sensitivity.
Simulations of PAUT inspections were performed using CIVA-UT to aid in the initial
procedure development. CIVA-UT is an analytical ultrasonic simulation software which can
compute ultrasonic beam properties and simulate ultrasonic inspection of various reflectors, including weld flaws. The modeling incorporated weld flaws that corresponded to the
critical flaw sizes developed previously. These flaws serve as a “lower bound” flaw set from
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
3 which an improved acceptance criteria were developed to consistently reject these flaws. As
long as flaws of this size or larger are consistently rejectable, the procedure will be effective at
removing critical flaws from service. Therefore, the acceptance criteria is grounded in fracture mechanics, although it does not use flaw height measurement for evaluation. Rather,
the future acceptance criteria primarily uses amplitude for flaw evaluation. The procedure
was verified through experimental testing of weld flaw specimens to verify that the detection
and rejection of critical weld flaws were improved.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
4
CHAPTER 1
Background
1.1 Problem Statement and
Research Objective
Reliable detection of internal weld flaws using any non­
destructive testing (NDT) technique is essential to ensuring the
desired performance of a structure. Presently, two NDT methods
are used for evaluation of complete joint penetration (CJP)
groove butt welds in steel bridges: radiographic testing (RT)
and ultrasonic testing (UT). Using RT, discontinuities are dis­
tinguished from sound weld or base metal based on contrast
variations that appear on a radiographic film. UT, on the other
hand, utilizes reflections from high-frequency sound waves to
inspect for internal discontinuities within the weld and base
metal. While RT can reliably identify volumetric discontinui­
ties, this method is typically not as effective for thin planar
discontinuities such as cracks and lack of fusion. UT typically
launches the sound waves at an angle into the material in such
a way that planar discontinuities can be readily detected.
Advances in ultrasonic methods, including the develop­
ment of phased array ultrasonic testing (PAUT), provide
enhanced ability to detect and characterize weld flaws, per­
form automated data collection, and generate images of
ultrasonic results. Although improvements have been made
to the ultrasonic equipment, the current acceptance criteria
for PAUT in the 2015 edition of American Welding Society
(AWS) D1.5 Bridge Welding Code [1] provided in Annex K
are not based on the criticality of a weld discontinuity on
bridge performance measures such as the resistance to fatigue
and fracture. Rather this acceptance criteria is a workman­
ship criteria meant to provide an arbitrary control on the
level of quality. In other words, the acceptance and rejection
criteria are not related to the structural performance of the
weld in any way. Because of the apparent “good experience”
with RT, thresholds for flaw rejection using conventional UT
were “calibrated,” though not systematically, to criteria used
traditionally for RT that also were not based on structural
performance [3, 4]. While this approach may seem reason­
able at first glance, it is absolutely critical to recognize that
RT and UT are totally different approaches for flaw detec­
tion simply due to the physics associated with the technolo­
gies. One technique responds to changes in density, which is
recorded on a 2-D film, while the other measures reflection
of sound in both the amplitude and time domain. Hence, the
common statement that “UT or PAUT can be used to see the
same flaws as RT” is incorrect.
Described herein is a summary of the proposed research
for the development of updated acceptance criteria for evalu­
ation of CJP welds in steel bridges using enhanced ultrasonic
testing methods. The proposed research included both ana­
lytical and experimental programs. The analytical testing
program included parametric studies of various plate geom­
etries, welds, residual stress fields, flaw types, and locations.
The analytical results were intended for establishing accep­
tance criteria and identifying the minimum flaw size that
must be detected reliably. The experimental testing program
had two components. The first consisted of developing pro­
cedures for characterizing and sizing flaws in CJP welds using
enhanced ultrasonic technologies. This included establishing
practical acceptance criteria that could be incorporated into
the AASHTO/AWS D1.5 Bridge Welding Code and guidelines
on their use. The second component consisted of a “round
robin” exercise in which plates with known flaws would be
circulated to a number of testing firms in order to evaluate
the proposed procedures and criteria.
The primary objectives of NCHRP Project 14-35 were to
use data collected through analytical studies and experimen­
tal testing to (1) develop guidelines to evaluate complete joint
penetration welds in steel bridges based on updated accep­
tance criteria and (2) develop proposed modifications to the
existing AASHTO/AWS D1.5 Bridge Welding Code.
As a minimum, the guidelines were to cover shop and field
fabrication and in-service evaluation and include procedures
for using enhanced ultrasonic testing methods to evaluate
CJP welds in steel bridges and pertinent acceptance criteria.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
5 1.2 Scope of Study
The research conducted for NCHRP Project 14-35, “Accep­
tance Criteria of Complete Joint Penetration Steel Bridge Welds
Evaluated Using Enhanced Ultrasonic Methods,” focused on
the application of PAUT to inspect steel bridge welds. Accep­
tance criteria, scanning procedure requirements, calibration
requirements, and technician qualification requirements were
developed using analytical and experimental methods. The
analytical methods included finite element models, fitness-forservice (FFS) models, and ultrasonic models. The experimental
methods included measurement of acoustic material properties
and ultrasonic testing of weld flaw specimens representing
typical bridge weld geometries. These specimens included
embedded and surface discontinuities representing both planar
and volumetric discontinuities. The ultrasonic testing of weld
flaw specimens included blind round robin testing by outside
conventional UT, PAUT, and TOFD technicians along with
inspection of the weld flaw specimens using RT. Weld flaw
specimens will also be used to perform final verification tests of
the proposed modifications to the PAUT inspection procedure.
1.3 Phased Array Ultrasonic
Testing (PAUT)
Since phased array relies on the same basic physics as con­
ventional UT to generate and receive ultrasound, many of the
details of PAUT inspection do not change from conventional
UT. However, unlike the single element transducer used in
conventional UT, PAUT uses multiple element transducers
and electronic time delays to generate and receive ultrasound.
Electronic time delay for
40° incidence angle
The electronic time delays use constructive and destruc­
tive interference that allow the ultrasonic beam to be
steered, scanned, swept, and focused electronically. Figure 1
shows the electronic time delays for 16 active elements of a
64-element transducer (i.e., elements 1–16 active) in order
to produce a 40° incidence angle (left) and 70° incidence
angle (right).
The array in the transducer can be constructed from a
linear array, a two-dimensional matrix array, or a circular
array. Linear arrays are used for most applications since
they are cheaper than more complex arrays and easier to
program [5]. Phased array probes commonly have between
16 to 128 elements. Focal flaws are calculated by the software,
which controls the time delays and firing sequence of the
transducer. The frequency of PAUT is very similar to conven­
tional UT, typically between 2–5 MHz for bridge weld testing.
Two types of scans are typically used for PAUT:
• Electronic scans (E-scans) are performed by multiplexing
the same focal flaw along a linear array. This will produce a
scan which is similar to manual scanned conventional UT.
• Sectorial scans (S-scans) are performed by altering the
time delays as the elements are fired, which creates a beam
which sweeps through a range of incidence angles.
PAUT can utilize encoded scanners to capture a continuous
stream of data from different transducer positions, either auto­
matically or semi-automatically. Semi-automatic scanning—
using a wheel or string encoder attached to the transducer—is
typically utilized for bridge welds due to the variation in
geometry associated with bridge fabrication. PAUTs using
Electronic time delay for
70° incidence angle
Figure 1. PAUT time delays.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
6
encoded scans have multiple views that can be displayed to
the technician, including the following:
• A-scan (x–y plot of amplitude vs. time for a single beam;
•
•
•
•
•
top left of Figure 2)
B-scan (end view when volume corrected)
C-scan (top view when volume corrected; bottom of
Figure 2)
D-scan (side view)
E-scan (end view of all A-scans when multiplexing same
focal law)
S-scan (end view of all A-scans for a range of incidence
angles, top right of Figure 2)
PAUT calibration involves correction of the wedge delay
and sensitivity calibration. For conventional UT, sensitivity
calibration involves measuring the reference amplitude of a
standard 1.5 mm diameter (0.06″) side-drilled hole (SDH)
reflector on an IIW-type calibration block; material attenu­
ation at other sound paths is accounted for by implement­
ing a correction through the attenuation factor equation. For
PAUT, the reference amplitude is calculated across the full
range of angles that will be used during the scanning. The
standard SDH reflector on the IIW-type block is still used,
but the beam is swept through all of the angles by moving the
transducer along the IIW-type block surface. After calibra­
tion, the reference reflector will have the same amplitude at
each angle (e.g., 70 degrees and 45 degrees). Time corrected
gain (TCG) is used to account for material attenuation by
sweeping the ultrasonic beam through SDH reflectors at
varying depths. After performing TCG calibration, identi­
cal reflectors will have the same amplitude regardless of the
depth or beam angle.
PAUT has many advantages over conventional UT, one of
which is the increased sound coverage. Compared to con­
ventional UT, PAUT can provide the UT technician with the
ability to scan a material using multiple beam angles simul­
taneously. The UT technician also has additional views such
as the S-scan and E-scan, which are two-dimensional repre­
sentations of all of the A-scans plotted simultaneously. This
can aid the technician in distinguishing false call signals due
to geometric indications. It can also help in flaw characteriza­
tion, through the use of tip diffraction signals or signals at the
surface. Weld overlays showing the geometry of the weld
preparation can also be drawn on the S-scan or E-scan views,
which can help UT technicians inspect locations where dis­
continuities are more likely, such as the fusion face or weld
root. If PAUT is used as a direct replacement of conventional
UT in manual raster scanning, these advantages are likely to
improve flaw detection and rejection if the same amplitudebased acceptance criteria were implemented.
Encoded PAUT scanning offers the ability to collect the raw
scan data and save it for future reference or viewing. Conven­
tional UT indications, on the other hand, are typically reported
Figure 2. Sample PAUT image (top left) A-scan, (bottom) C-scan, and (top right) S-scan.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
7 as tabulated values of indication amplitude, length, and loca­
tion. Operator error is introduced into the reporting process
since these values often are manually transferred from physical
measurements or instrument results. Conventional UT A-scan
data is also not typically saved for future reference.
Although PAUT can provide more coverage than conven­
tional UT, full coverage of the weld does not ensure that all dis­
continuities within the covered region will be detected. When
line scanning is performed with a single transducer, each point
in the volume of the weld will only be primarily covered by
sound with a single angle of incidence (it is recognized that
due to beam spread, a given location will be “hit” by more than
one angle of incidence but not with significant amplitude). If
the flaw is not oriented in a manner to reflect adequate ultra­
sound back to the transducer based on the specific angle of
incidence, the discontinuity may not be detected (or very little
sound reflected) even though sound is covering that region.
For this reason, it is often recommended to scan with angles
that are normal to “expected” discontinuities, such as fusion
faces of welds.
When line scanning is performed, the probe is typically
kept normal to the weld axis to inspect for discontinuities
which are primarily oriented parallel to the weld axis. Con­
ventional UT, on the other hand, is typically performed by
raster scanning where the probe is moved with rotation,
transverse, and longitudinal movements. This movement
helps to maximize the amplitude response from discontinu­
ities that are not oriented perfectly parallel to the weld axis.
Prior PAUT research found that a skew angle of only 10°
from the alignment of the discontinuity caused the signal
amplitude to drop considerably and flaw detection become
marginal [6]. A change in skew angle of 20° from perpen­
dicular to the discontinuity resulted in total loss of disconti­
nuity response. Therefore, lack of raster scanning when line
scanning with PAUT is likely to result in decreased ampli­
tude for some weld flaws.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
8
CHAPTER 2
Research Approach
2.1 Summary of the State
of the Practice
2.1.1 Current AWS D1.5 Requirements
A set of acceptance criteria provides a measure or reference by which a standard of quality is applied to provide adequate structural performance. One definition of acceptance
criteria is “a set of rules formulated in terms of the requirements to NDE recorded parameter values for judgement of
whether flaws are acceptable or rejectable [7].” It would be
ideal for acceptance criteria to reject and repair all imperfections which could be harmful to the structure while accepting
all harmless imperfections, but this idea is unattainable in a
rational weld acceptance criteria. If one takes into account
the economic considerations of repairing a fatigue failure
due to undersizing a flaw compared to the cost of repairing
a benign imperfection, it may be found that the acceptance
criteria needs to be set so that many harmless imperfections
may need to be repaired in order to eliminate one harmful
imperfection [8].
On a very broad level, all acceptance criteria can be placed
into one of two categories: workmanship criteria or fitnessfor-service (fracture-mechanics-based) criteria. Workmanship criteria are based on a general, arbitrary control on the
level of quality [3] and is aimed at ensuring that an acceptable
workmanship level is met [7]. Many welding codes employ a
workmanship criteria, including AWS D1.1 Structural Welding Code [9] and AWS D1.5 Bridge Welding Code. Generally
speaking, although workmanship criteria have historically
provided adequate performance, they are often based on
experience and do not give an objective comparison to the
actual size that would result in component failure. Further, the
apparent “success” of workmanship criteria (i.e., the observation that no problems mean that the criteria are working)
may not necessarily be due to the criteria themselves, but due
to a series of factors that are unknown or unaccounted for
since the criteria were arbitrarily crafted.
Fitness-for-service (FFS) criteria, also known as Engineering Critical Assessment (ECA), are based on fracture mechanics which uses information on member loading and material
properties to determine an acceptable initial discontinuity size
for the intended service life. FFS will typically permit larger
discontinuities than workmanship criteria but require accurate and reliable measurements of flaw size and location [10].
Further, FFS requires accurate estimates of material properties
and residual stresses, in addition to static and cyclic stresses
over the service life of the structure.
AWS D1.5 conventional UT employs workmanship criteria based on the amplitude of the reflected sound along with
the flaw length. Conventional UT technicians perform bridge
weld testing under the AWS D1.5 code by utilizing a manual
raster scanning approach where the probe is rotated and translated on the testing surface to provide coverage of the entire
weld volume and to maximize the signal response amplitude.
In AWS D1.5, thresholds for flaw rejection using conventional
UT were developed through calibration to criteria used traditionally for RT that were not based on structural performance
[3, 4]. RT and UT utilize very different approaches for discontinuity detection due to the actual physics associated with the
technologies. For example, RT responds to changes in density,
which are recorded on a 2-D film, while UT measures reflection of sound in both the amplitude and time domain. It
cannot be assumed that UT or PAUT can always detect the
same discontinuities as RT. Generally, it has been found that
UT methods have increased sensitivity to crack-like flaws
while RT has increased sensitivity to volumetric flaws.
For conventional UT inspections, according to AWS D1.5,
the indication rating is determined based on the indication
amplitude compared to the reference standard reflector
and the sound path distance. Decreasing values (i.e., more
negative values) of indication rating are more severe. The
indication rating is derived by subtracting the reference gain
and an attenuation factor from the equipment gain when the
indication amplitude matches the reference amplitude. The
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
9 attenuation factor is included to account for the ultra­
sonic attenuation due to the loss of amplitude as the sound
travels through the steel. Discontinuities larger than the
reference reflector, which is a 1.5 mm (0.06″) diameter
SDH, should reflect more sound than the reference reflector. Therefore, the equipment gain will be lower when the
indication amplitude matches the reference level. A negative indication rating would result if the sound traveled
the same distance in the inspection as the reference reflector
and lower equipment gain is required to match the reference
amplitude. Assuming that the sound path remains the same,
a positive indication rating would result if more equipment
gain is required to match the reference amplitude.
Based on the loading (compression or tension), indication
rating, plate thickness, and testing angle, the indication is
classified by assumed severity:
• Class A (large flaws): Any indication in this category is
rejected (regardless of length).
• Class B (medium flaws): Any indication with a length
greater than ¾ inch is rejected.
• Class C (small flaws): Any indication in this category
with a length greater than 2 inches or ¾ inch for an indication in the top or bottom quarter of a tension weld is
rejected.
• Class D (minor flaws): Any indication in this category is
accepted regardless of length or location in the weld.
For plate thicknesses up to 1.5 inches, the range for
intermediate classifications (i.e., Class B and Class C) is
1 decibel (dB). Therefore, only 3 dB separates a Class A
(automatically rejectable) indication from a Class D (automatically acceptable) indication. For plate thicknesses
greater than 1.5 inches, the range for intermediate classifications is 2 dB, and 5 dB separates a Class A indication from
a Class D indication.
AWS D1.5:2015 includes alternate acceptance criteria in
Annex K to allow for the implementation of PAUT in lieu
of conventional UT for testing of bridge welds. This testing procedure employs a line scanning approach where the
probe remains perpendicular to the weld at a constant index
position. The procedure uses a sectorial focal law which produces a sound wave over a range of incidence angles. This
helps to insonify the weld volume. However, multiple scans
at varying index points may be necessary for complete coverage. The acceptance criteria in Annex K were developed
as an adaptation of an existing conventional UT acceptance
criteria in AWS D1.1 (Annex Q) and were also workmanship criteria; amplitude of the reflected sound along with the
flaw length form the basis of the acceptance criteria. Similarly, Annex K uses the same size reference standard reflector
and the same indication classifications (Class A–D). As will
be discussed further in Chapter 3, the acceptance criteria
in AWS D1.1 Annex Q and subsequently AWS D1.5 Annex K
do not match the acceptance criteria in AWS D1.1 or
D1.5 Clause 6 for conventional UT. While the classifications
and their respective maximum length requirements are very
similar for Annex K and conventional UT, the range in intermediate classifications (i.e., Class B and Class C) are much
larger. Class B has a 5 dB range, and Class C has a 6 dB
range. Therefore, 11 dB separates a Class A (automatically
rejectable) indication from a Class D (automatically acceptable) indication. For Annex K, the reference amplitude is
consistently used as the distinction between a Class B or
Class C indication, while the reference amplitude does not
correlate to a distinct flaw classification in the conventional
UT tables.
Instead of using an indication rating and an attenuation factor to evaluate the amplitude of the indication such
as is performed in Clause 6 conventional UT inspections,
PAUT utilizes a calibration method with reference reflectors
placed at various depths (i.e., TCG). With this correction,
the amplitude measured in percentage of full screen height
(%FSH) is compared directly with the reference amplitude.
Indications with a greater amplitude in %FSH, therefore,
are more severe, unlike conventional UT where more negative indication ratings are more severe.
2.1.2 Comparison to Other UT Codes
A collection of reference standard provisions, both
national and international, related to ultrasonic testing have
been summarized below. Specifically, a comparison of each
standard’s policy on acceptance criteria, material attenuation,
and probe frequency has been presented in Table 1, Table 2,
and Table 3, respectively. The ultrasonic codes included in this
summary are as follows:
• Canadian Standards Association (CSA) W59 code, appli-
cable to bridges [11]
• European Standard (EN) and International Organization
for Standardization (ISO) codes, applicable to bridges
• American Society of Mechanical Engineers (ASME) Boiler
and Pressure Vessel Code (BPVC), applicable to the gas and
nuclear industry [12]
• Japanese Industrial Standard (JIS) Z 3060 code, applicable
to bridges [13]
2.1.2.1 Acceptance Criteria
Codes that included an acceptance criteria based on measuring the flaw size of weld flaws require that the flaw sizing
procedures be developed by the PAUT technician and verified
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
10
Table 1. Acceptance criteria summary.
Specification
Acceptance
criteria based on
flaw sizing using
PAUT
Prescribes flaw
sizing procedure
AWS D1.5-15
X
n/a
CSA W59-18
Requires
performance
qualification
X
X
*
ISO 19285-17
ASME BPVC
CC 2235-13
JIS Z 3060-15
Acceptance criteria
based on max
amplitude & length
using PAUT
(Flaw Sizing)
X
X
X
X
X
X
X
*CSA W59-18 does not provide any specific acceptance criteria for PAUT based on flaw sizing but gives minimum
requirements and allows for other acceptance criteria which have been deemed to be equivalent.
Table 2. Material attenuation summary.
Specification
AWS D1.5-15
Accounts for the material
attenuation due to varying
grade or microstructure
No
CSA W59-18
No (UT) / yes(PAUT)
ISO 17640-17
Yes
ASME
BPVC-17
Yes
JIS Z 3060-15
Yes
If so, how?
Conventional UT
PAUT
n/a
n/a
Qualification testing through
n/a
same medium.
Requires a calibration block. If the calibration block
and test object are not acoustically the same, a transfer
correction is to be applied.
Requires a calibration block of the same product form
and material specification of the material being
examined. If any acoustic differences remain between
the calibration block and test object, a transfer
correction is to be applied.
Requires all calibration blocks to be of a steel material
with equivalent acoustic characteristics to the test
object.
Table 3. Probe frequency summary.
Specification
AWS D1.5-15
CSA W59-18
Probe Shear Wave Frequency Range
(MHz)
Conventional UT
PAUT
2–2.5
1–6
2–2.5 (Fixed
Attenuation)
2.25–10 (TCG)
2.25–10 (TCG)
Notes
—
—
ISO 17640-17
2–5
No stipulation
Lower frequencies may be necessary for
testing at long sound paths and/or high
material attenuation
ASME
BPVC-17
1–5
1–5
—
JIS Z 3060-15
2–5
2–5
Sound path length stipulations are put on
using higher frequencies
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
11 for accuracy through performance qualification of the PAUT
procedure on weld mockups representative of those being
inspected before performing testing. None of the PAUT codes
provide a prescriptive procedure for measuring the throughthickness height of weld flaws.
AWS D1.5 Annex K uses maximum amplitude and length
for the acceptance criteria for PAUT. It does not require that
any performance qualification testing be performed. It outlines the requirements of performing a mockup verification
at the option of the PAUT technician or when required by
the engineer.
CSA W59 added a TCG approach in the 2018 edition for
conventional UT or raster scanned manual PAUT, which is
intended to provide an equivalent level of quality as the current conventional UT acceptance tables but has one table
for all angles. These changes were compared to AWS D1.5
Annex K, and it was found that the CSA W59 acceptance criteria will generally be conservative compared with the current
AWS method [14–16]. The CSA code also provides requirements for use of encoded line scanned PAUT or other alternative ultrasonic systems in lieu of conventional UT if agreed to
in writing by the engineer and contractor prior to inspection.
In order to use encoded PAUT, it requires that a written procedure be developed and that performance qualification tests
of the procedure be performed to verify that the minimum
required sensitivity is provided. No prescriptive procedures
are provided for flaw sizing or scanning of the welds.
ISO 19285:2017 [17] provides acceptance criteria for PAUT
which may be applied to bridge welds. This code allows for
either evaluating the welds using the flaw size (e.g., flaw height
and length) or maximum amplitude and flaw length. This
code requires performance qualification for all PAUT inspection procedures on a test block of the same material and similar thickness as the test object with reflectors of prescribed
size and location. No prescriptive procedures are provided
for flaw sizing, and verification of flaw sizing procedures is
required.
ASME BPVC Code Case 2235-13 [18] provides acceptance
criteria for PAUT in lieu of RT for the nuclear and petrochemical industries. The code case allows that evaluation of
final acceptance only be performed by flaw sizing, but amplitude may be used for detection. The code case requires performance qualification of all PAUT inspection procedures
on a test block of the same material with multiple reflectors
throughout the thickness of the part. Requirements are given
for the size and location of the reflectors. No prescriptive procedures are provided for flaw sizing, and verification of flaw
sizing procedures is required.
JIS Z 3060-2015 does not provide an acceptance criteria
specifically for PAUT. This code provides classification of
discontinuities based on conventional UT results, but no
acceptance criteria is included.
2.1.2.2 Material Attenuation
AWS D1.5 Annex K is the only code that does not require
PAUT technicians to account for differences in material
attenuation between the calibration block and the test specimen. In fact, there is no discussion on the acoustic properties of the calibration block compared to the test object in
the AWS code. For conventional UT, both the AWS D1.5 and
CSA W59 account for attenuation in the test specimen by the
application of an attenuation factor and use of an International Institute of Welding (IIW)-type reference block to set
the reference amplitude. However, the 2018 edition of CSA
W59 provides a TCG approach in lieu of the fixed attenuation approach for conventional UT, while the TCG approach
is required for manual raster scanned PAUT. While there is
limited discussion in the CSA code on how to account for
differences in material attenuation, it does state that the calibration block should be “acoustically equivalent” to the test
object. For encoded line-scanned PAUT inspections, the CSA
code requires a calibration procedure be developed on a caseby-case basis as part of a written procedure.
The ISO and ASME codes specifically state that modifications to calibration are required if the material attenuation differs between the calibration block and the test object,
including both base metal and weld metal. This is typically in
the form of a transfer correction. Methods for determining
the transfer correction are described in detail in Section 3.4.2.
ISO 17640 [19] requires a transfer correction be applied
when a difference of 2 dB to 12 dB is observed at the longest
inspection sound path. Any difference less than 2 dB is
negligible, and any difference greater than 12 dB is a cause
for reevaluation of the calibration procedures. ASME BPVC
[12] states that if “the block material is not of the same
product form or has not received the same heat treatment, it
may be used provided it meets all other block requirements
and a transfer correction for acoustical property differences
is used.” ASME does not provide requirements on the use
of a transfer correction; instead, it is left to an inspector’s
discretion. JIS Z 3060 [13] provides five different calibration
blocks to be used in different circumstances. Each reference
block is required to be of a steel material with equivalent
acoustic characteristics to the test object.
A single, unanimous definition of the phrases acoustically equivalent, acoustic characteristics, or acoustic properties
does not exist. ISO [20] defines acoustical properties as the
characteristics of a material which control the propagation
of sound in a material. In ultrasonic testing, these principal
characteristics are ultrasonic velocity and attenuation. For
example, ASTM E114 [21] states that “the reference standard
material and production material must be acoustically similar (in velocity and attenuation).” JIS Z 3060 states that the
difference in the ultrasonic velocity of the test object and the
calibration block shall be within ±2% and that the transfer
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
12
correction shall be within ±2 dB. Measurement of the acoustic properties requires ultrasonic testing of the test object and
calibration block including the use of normal incidence shear
wave probes and/or pitch-catch methods. ISO 2400, the standard for the IIW calibration block, imposes strict requirements on the material, heat treatment, and surface finish of
the IIW reference block. Following fabrication with these
guidelines, the acoustic velocity of the block must be checked
and fall within ±0.2% of the prescribed wave velocities.
Ultrasonic velocity and attenuation may be the two material
characteristics to have the biggest impact on ultrasonic evaluation, but grain size, grain structure, material composition,
and surface roughness are all factors determining the velocity
and attenuation of a material.
2.1.2.3 Probe Frequency
AWS D1.5 and CSA W59 restrict the probe frequency for
conventional UT due to the fixed attenuation factor, which is
only valid for a specific probe size, shape, and frequency [15].
The 2018 edition of the CSA W59 code allows a wide range
of probe frequencies for the TCG approach, but, as noted
above, this requires that the calibration block be acoustically
equivalent to the test object. The other codes allow a wider
range of probe frequencies but also require that a calibration
be performed to take into account material attenuation. The
ISO 17640 code [19] states that lower frequencies are recommended for conventional UT where acceptance is determined based on maximum amplitude and length rather
than flaw characterization and sizing. While removed for
the 2017 edition, the 2010 edition of ISO 17640 stated that
initial testing use frequencies as low as possible, but within
the specified range. JIS Z 3060 [13] stipulates the allowed
probe frequency be determined based on the sound path
distance with longer sound paths having lower frequencies. JIS Z 3060 allows 3.5–5 MHz probes be used on sound
paths that are 100 mm (3.9″) or less. For second leg scans,
this limit would be exceeded for a thickness of 1.4″ at a 45°
incidence angle and at a thickness of 0.7″ at a 70° incidence
angle. Anything over 250 mm (9.8″) is only allowed to be
inspected using 2 MHz.
2.2 Research Methodology
Four major research efforts constitute the approach taken
to attain the objectives of the current research. The first was
the evaluation of the critical flaw size for CJP bridge welds
based on fracture mechanics using FFS methodology. Once
the critical flaw size was determined, an experimental round
robin testing program was then conducted using weld flaw
specimens to compare test results using PAUT in accordance
with AWS D1.5 Annex K to conventional UT, TOFD, and RT
and to aid in the development of improvements to Annex K.
The third step involved experimental testing that evaluated
the acoustic properties of bridge weld and base metals and the
development of calibration requirements. The fourth and
final effort involved the use of computer modeling of ultrasonic testing along with experimental ultrasonic testing of
weld specimens in developing recommendations for revised
scanning requirements and acceptance criteria in detecting
and rejecting critical weld flaws.
2.2.1 Evaluation of Critical Flaw Size
for Steel Bridge Welds
An analytical parametric study was performed to establish
the critical flaw size that would be considered rejectable for
typical bridge CJP welds. The results were utilized in identifying the critical flaw size that should be reliably detected
and rejected and to establish revised acceptance and rejection criteria. The critical flaw sizes were evaluated based on
fracture mechanics using a BS 7910:2013 [2] FFS approach.
These internal flaws are assumed to be cracks. The critical
size of volumetric flaws was also evaluated through review
of prior fatigue test results. This review highlighted that the
fatigue crack growth threshold of volumetric flaws is similar
to that of planar flaws. Therefore, FFS analysis can be applied
for volumetric flaws by evaluating the maximum flaw size
that will not undergo fatigue crack growth.
FFS was performed to evaluate two different failure modes
for planar flaws. The first is failure due to fatigue crack
growth to ensure that internal defects do not grow in under
cyclic loading. The second is failure due to brittle fracture.
The analytical studies that were performed under this task
included parametric studies of various plate geometries,
welds, residual stress fields, flaw types, flaw sizes, and locations. The results from the analytical program have identified
the target critical flaw size.
2.2.2 Round Robin Experimental
Testing Program
A round robin testing program was performed in order to
gain insight into the capabilities of the current technicians
in the steel bridge industry and to identify best practices
for improved flaw detection and flaw characterization. The
round robin testing program was used to determine the minimum flaw size that could be reliably detected with enhanced
ultrasonic methods and how the advanced methods compare
with the historical conventional UT method. The round robin
experimental testing program was performed by circulating
weld flaw specimens among technicians in order to acquire
inspection data from PAUT, conventional UT, TOFD, and RT.
This data was used to compare the inspection results from
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
13 the various inspection methods. This comparison included
determination of the hit/miss rate, rejection rate, frequency
of false calls, accuracy of flaw height and length sizing, accuracy of flaw type characterization, and technician variability.
The results of the round robin testing program were used
to improve development of future PAUT scanning procedures and acceptance criteria. These results highlighted that
moving forward with the development of acceptance criteria
based solely on the measurement of the flaw size is not
feasible at present with the inaccuracy and unreliability in
flaw size measurements using the current PAUT workforce.
Therefore, the focus of the research was on improving the
acceptance criteria based on maximum amplitude and flaw
length in AWS D1.5 Annex K. An acceptance criteria based on
physical measurements of the flaw height and length has also
been provided in order to allow for alternative UT inspection
methods or future technological improvements.
2.2.3 Development of Calibration
Requirements for Variations
in Acoustic Properties
Experimental testing of various bridge welds and base
materials has highlighted that the acoustic properties may
vary considerably, including acoustic attenuation and shear
wave velocity. Additional calibration requirements are necessary in order to account for possible differences between the
calibration block and the test object. Not properly account-
ing for these acoustic properties could result in large variations of the amplitude of indications in the test object.
The experimental results of the acoustic properties of
various bridge base materials were used to generate benchmarked material models for ultrasonic inspection simulations using CIVA-UT. These simulations were used to
develop requirements on the calibration block material and
scanning parameters in order to limit possible error in reference sensitivity.
2.2.4 Development of PAUT Scanning
Procedures and Acceptance Criteria
Simulations of PAUT inspections were performed using
CIVA-UT to aid in the initial procedure development. The
modeling incorporated weld flaws corresponding to the
critical flaw sizes previously developed. These flaws serve as
a “lower-bound” flaw set from which improved acceptance
criteria were developed to consistently reject these flaws. As
long as flaws of this size or larger are consistently rejectable,
the procedure will be effective at removing critical flaws from
service. Thus, the proposed acceptance criteria are grounded
in fracture mechanics although they do not explicitly use flaw
height measurement for evaluation. Rather, the future acceptance criteria primarily use amplitude for flaw evaluation.
The procedure was verified through experimental testing of
weld flaw specimens to verify that the detection and rejection
of critical weld flaws were improved.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
14
CHAPTER 3
Findings and Application
3.1 Critical Flaw Size
Analytical parametric studies were performed in order
to establish the critical flaw sizes considered rejectable for
typical bridge CJP welds. The results were used to identify the
critical flaw size that must be reliably detected and rejected
to establish revised acceptance and rejection criteria. The
critical flaw sizes were evaluated based on fracture mechanics
using a BS 7910 [2] FFS approach, and the internal flaws are
assumed to be cracks. The parametric studies that were performed included various plate geometries, welds, residual
stress fields, flaw types, flaw sizes, and locations.
3.1.1 Volumetric Flaws
Flaw severity is dependent on the flaw type (i.e., planar vs.
volumetric). Therefore, volumetric (i.e., non-planar such as
slag and porosity) flaws and planar flaws were considered separately. Fracture mechanics relies on an underlying assumption that a flaw is a crack, whether considering fatigue crack
growth or fracture. This is a very conservative assumption
for smooth, rounded volumetric flaws, but assuming that all
volumetric flaws are void of any cracks or crack-like geometry is an unconservative assumption to make. During the
literature review, it was apparent that all existing enhanced
ultrasonic testing specifications do not distinguish between
flaws that are volumetric as opposed to crack-like, other than
including additional criteria for rejecting scattered volumetric flaws, such as a limit on the number of point-like reflectors over a specific weld length.
In the round robin testing performed during this research
project, inspectors were requested to document whether
a detected weld discontinuity was volumetric or planar.
Responses showed that volumetric discontinuities and planar
discontinuities could not be differentiated with reasonable
accuracy. Not only were volumetric discontinuities often
reported as being planar, but planar discontinuities were
often reported as being volumetric. It was clear that the evaluation of truly planar discontinuities using an acceptance
criteria developed for volumetric discontinuities will not
capture the criticality of planar discontinuities and would
be unconservative. Therefore, it was determined that PAUT
acceptance criteria should not require flaw characterization
as planar and volumetric flaws, but rather all indications
should be evaluated against the same acceptance criteria.
In order to develop target critical flaw size for volumetric flaws, existing experimental test data on CJP welds with
such flaws was explored. Harrison [22–24] investigated the
fatigue resistance and strength of butt welds with embedded slag inclusions and porosity. That research found that the
effect of slag inclusions and porosity below 10% volume on
ductile strength was negligible due to the overstrength of the
weld metal. The porosity limit was set to 10% since this was
considered the limit that should be allowed without masking
other indications during RT inspection. Slag inclusions were
found to typically have limited through-thickness height—
the critical dimension according to fracture mechanics—due
to occurring between weld passes. During these studies, data
from high-cycle fatigue tests performed by multiple researchers were collected by Harrison to develop an acceptance criteria
based on S-N curves. This was used to set limits on the percentage of the volume of weld metal containing porosity and the
maximum length of slag inclusions that were later incorporated
into BS 7910. These limits assume that the flaw is verified to
be volumetric and was developed to be used with RT. BS 7910
states that “the tolerable porosity sizes based on ultrasonic testing may be considerably less, particularly for thinner sections.”
This is because BS 7910 recognizes that volumetric flaws can be
easily undersized by ultrasonic testing. BS 7910 is silent on how
specifically the allowable sizes should be reduced for ultrasonic
testing; however, this would likely be determined through some
kind of performance testing and POD evaluation.
The critical dimensions for slag inclusions in BS 7910:2013
are given in Table 4 when combining the height requirements
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
15 Table 4. Combined BS 7910 slag fatigue and fracture
requirements.
Stress Range (ksi)
6.09
5.37
4.64
4.06
3.34
Slag Height (2a)
(0.5" thick weld)
0.063"
0.063"
0.063"
0.063"
0.063"
for fracture (no length requirement for fracture) and the
length limit for fatigue (no height requirement for fatigue)
assuming infinite fatigue life. As shown in this table, the critical dimensions for slag inclusions in BS 7910 were already
quite small before considering the effect of undersizing with
ultrasonic testing.
AWS D1.5 RT acceptance criteria include limits on maximum indication length and spacing. While these limits are
workmanship-based and not based on experimental results,
they do help by highlighting the typical flaw size that has been
traditionally accepted. A minimum slag height for RT sensitivity was also calculated by comparing the size and density of
required wire-type image quality indicators (IQI) to determine
the required slag height for the same change in density. This
analysis assumed the density of slag to be 3.5 g/cm3 based on
literature [25, 26]. The maximum slag length and minimum
slag height for each plate thickness are shown in Table 5. This
confirmed that the BS 7910 maximum slag height would likely
be detectable under RT inspection and that the BS 7910 maximum slag length requirements compare reasonably to the AWS
D1.5 RT acceptance criteria for stress ranges of ∼5 ksi.
The critical planar flaw sizes must also be considered when
evaluating critical volumetric flaws since calculation of the
stress intensity factors (KI) shows that any sharp or planar
flaw extending from a slag inclusion or porosity will result in
an equivalent planar flaw extending over the combined area
of the volumetric and planar flaws. Thus, once a crack begins
to extend from a volumetric flaw, it is equivalent to a crack
which extends over the entire projected area.
An appropriate delta-K threshold (DKth) to control fatigue
crack growth of volumetric flaws was evaluated using fatigue
test data. Using the combined fatigue and fracture limits on
slag given by BS 7910 in Table 4, it was found that the DKth
Table 5. Slag RT sensitivity and AWS D1.5 length
requirements.
Plate Thickness (in ")
0.5"
1"
2"
3–4"
Min. Slag Height (2a)
0.028"
0.044"
0.057"
0.089"
Max. Slag Length (2c)
0.125"
0.313"
0.500"
0.500"
Slag Height (2a)
(> 1" thick weld)
0.118"
0.118"
0.118"
0.118"
0.118"
Slag Length
(2c)
0.098"
0.157"
0.394"
1.378"
Infinite
varied from 1.5 to 2.0 ksi in depending on the stress range.
In addition, NCHRP Report 335: Acceptance Criteria for Steel
Bridge Welds [8] recommended using 2.5 ksi in for DKth in
the development of a modified RT acceptance criteria for
bridge welds. NCHRP Report 335 included fatigue testing
on weld samples with porosity and slag weld imperfections
and reported that the calculated initial DK ranged from 2.25
to 3.7 ksi in . As will be discussed in the following section,
2.5 ksi in was used for DKth in this project for determining
the critical flaw size for planar flaws. Since this is also appropriate for volumetric flaws, planar and volumetric flaws have
the same critical flaw size based on fatigue. Therefore, the target critical planar flaw sizes shown in the following sections
were used to develop target critical volumetric flaw sizes for
the CIVA parametric matrix.
3.1.2 Critical Planar Flaw Sizes
Based on Fatigue
Various planar flaws were modeled using FFS procedures to
determine the maximum flaw size that will not grow in fatigue
(i.e., infinite life). Failure was defined as the limiting flaw size
at which crack growth is expected at a given stress range since
the intention is to evaluate welds in a fabrication shop environment where the anticipated annual daily truck traffic (ADTT)
would not need to be taken into consideration for each individual weld. The inputs for the analysis accounted for variations in the plate thickness, flaw aspect ratio, flaw position,
and magnitude of the applied live load stress range. Thickness
transition weld geometries were also considered by accounting
for stress concentrations through finite element analysis. It is
noted that one could utilize a finite life approach and assume a
number of cycles per day over the design life which could result
in larger flaw tolerance; however, such an approach was not
deemed to be implementable for several reasons. For example,
the designer would need to provide detailed static and cyclic
stress data for each weld. Then, the inspector would need to
use this data to select the appropriate inspection criteria for
the specific weld. Clearly, there is much room for error in this
approach in addition to the fact it places considerable responsibility on the technician to interpret the data.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
16
3.1.2.1 Cyclic Stress Range
To determine the stress range for evaluation, a reasonable
approach is to use the stress range associated with infinite life;
for example, using a stress range of 16 ksi for a Category B butt
weld detail. However, the stress range associated with the CAFL
(constant amplitude fatigue limit) of Category B (16 ksi) is very
high in terms of actual in-service stress ranges for a fatigueresistant detail such as a CJP butt weld. The FFS calculations
indicated that a 16 ksi stress range results in a very small maximum permissible flaw size to ensure infinite life. It is noted
that NCHRP Report 335 recommends using about half of the
fatigue limit for the detail under consideration, for example,
8 ksi to 9 ksi for Category B, as it better represents the upper
bound in-situ effective stress range. NCHRP 335 even suggests using lower values if data exists for justification or if
deemed acceptable by the engineer. Therefore, critical crack
sizes were calculated for stress ranges from 4 ksi to 20 ksi in 4
ksi increments during the parametric study.
3.1.2.2 Delta-K Threshold (DKth )
The delta-K threshold (DKth), which correlates to initiation
of fatigue crack growth, was determined through review of
previous research that had performed fatigue testing of welds
in structural steels, including NCHRP Report 267: Steel Bridge
Members Under Variable Amplitude Long Life Fatigue Loading [27] and NCHRP Report 181: Subcritical Crack Growth
and Fracture of Bridge Steel [28]. The appropriate value of
DKth depends on the R ratio which is the ratio of minimum
applied stress to maximum applied stress. The typical R ratio
at a butt weld is quite high due to the presence of residual
and dead load stresses. According to Barsom and Novak [28],
DKth is equal to 2.05 ksi in for an R ratio of 0.8. According to
Fisher et al. [27], the constant amplitude crack growth threshold approaches 3 ksi in at high R ratios, but he found that
DKth approaches 2 ksi in for an R ratio of 0.8 for random variable block loading when only a small percentage of stress
intensity factor ranges exceeds the constant amplitude crack
growth threshold.
Therefore, based on the previous testing performed on
DKth for typical structural steels, 2.5 ksi in was used for DKth
in this project for both planar and volumetric flaws (i.e.,
planar and volumetric flaws have the same critical flaw size
for fatigue). While the specific residual stresses were not
considered in the fatigue evaluation, the effect of the residual
stresses on the fatigue crack growth was accounted for in
the determination of the DKth through the R ratio. Since
infinite life was utilized in the analysis, the effect of residual
stresses on the da/DN curves did not affect the analysis.
Therefore, the yield strength (i.e., grade) is not a variable in
the test matrix.
3.1.2.3 Summary of Parametric Study
Using the inputs of the fatigue-loading parametric study,
the maximum crack height and length that would initiate
crack growth for CJP butt welds was determined for various
a/c ratios, crack position, plate thickness, stress range, and
type of weld (equal thickness or thickness transition). Figure 3 shows typical crack position and weld type. Signal FFS
software [29] was used to perform this analysis by converting
the fatigue crack growth problem to an equivalent fracture
problem by setting the fracture toughness (Kc) equal to the
DKth value of 2.5 ksi in and the primary stress equal to the
applied stress range. The specific joint types that were evaluated for fatigue are summarized in Table 6.
The fatigue analysis and findings are presented in the following sections, along with the fracture analysis findings for
each geometry type including equal thickness butt welds and
thickness transition butt welds. The results of T and corner
joints are presented in Appendix A.
3.1.3 Critical Planar Flaw Sizes
Based on Fracture
Similar to the fatigue loading models, planar flaws have
been modeled using FFS procedures to determine the maximum crack size that will not result in fracture. The Option 1
FFS procedure found in BS 7910:2013 was utilized to evaluate
Figure 3. Typical flaw positions within weld.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
17 Table 6. FFS evaluation matrix for CJP butt welds.
Type of
Crack/Position
Applied Stress
Range
(Category B)
4 ksi, 8 ksi, 12 ksi,
16 ksi, 20 ksi
Type of Butt
Weld
A/C Ratio
Thickness
Equal
Thickness
0.01-1.5
0.5", 1", 2", 3", 4"
Transition
0.01-1.5
0.5" to 1", 1" to 1.5", 1"
to 2", 2" to 3", 2" to 4",
3" to 4"
4 ksi, 8 ksi, 12 ksi,
16 ksi, 20 ksi
Equal
Thickness
0.01-1.5
0.5", 1", 2", 3", 4"
4 ksi, 8 ksi, 12 ksi,
16 ksi, 20 ksi
Transition
0.01-1.5
0.5" to 1", 1" to 1.5", 1"
to 2", 2" to 3", 2" to 4",
3" to 4"
4 ksi, 8 ksi, 12 ksi,
16 ksi, 20 ksi
Equal
Thickness
0.01-1.5
0.5", 1", 2", 3", 4"
4 ksi, 8 ksi, 12 ksi,
16 ksi, 20 ksi
Transition
0.01-1.5
0.5" to 1", 1" to 1.5", 1"
to 2", 2" to 3", 2" to 4",
3" to 4"
4 ksi, 8 ksi, 12 ksi,
16 ksi, 20 ksi
Near Surface
¼ depth
Mid-depth
the effects of various flaw types on the performance of CJP
welds. While finite element models were developed to evaluate the effects of plasticity for transition welds, the Option 1
failure assessment diagram (FAD) curves were used to account
for the inherent variability in the fracture mechanics inputs.
A parametric study was performed to account for variations
in the plate thickness, flaw aspect ratio, flaw position, magnitude of the combined dead and live stresses, residual stresses,
and material properties (i.e., strength and fracture toughness).
3.1.3.1 Yield and Tensile Strength
Both yield and tensile strength must be included in the
fracture analysis. Since steel products are generally delivered
above the minimum specified yield strength, the analyses
assumed that the expected yield strength is 10% greater
than the minimum specified yield strength (i.e., FyExpected =
1.1FyNominal). Similarly, the expected tensile strength was also
assumed to be 10% greater than the minimum specified tensile strength. The expected yield and tensile strength of the
as-placed weld were assumed to match the expected yield and
tensile strength of the base material. Specific material grades
considered in the parametric study were 36, 50, 70, and 100 ksi.
3.1.3.2 Fracture Toughness
A fracture assessment requires toughness to be known or
at least assumed. At present, the U.S. bridge industry does
not require any specific level of toughness or testing for
as-placed production CJP welds or for the heat-affected zone
(HAZ). In the absence of any specifications, one could conservatively use the lower-bound estimate of KIc per the master
curve of around 20 ksi in , where KIc is the critical fracture
toughness at initiation of fracture. However, this will result
in extremely conservative estimates of tolerable flaw sizes.
Rather than assume a lower-bound toughness, the material
toughness was included in the fracture analyses as a variable
in the parametric study. Toughness was varied from a lower
bound of 25 ksi in to 100 ksi in in 25 ksi in increments
to establish the sensitivity and relationship between flaw size
and toughness. It was assumed that the toughness used in the
model corresponds to that of the as-placed weld and HAZ at
the lowest anticipated service temperature.
Charpy v-notch (CVN) requirements are included in
ASTM A709 [30] for base metal of production welds and
AASHTO/AWS D1.5 [1] for weld metal of groove weld
qualification test plates. These welds do not necessarily
match the thickness or joint geometry of the production
welds. While these requirements were used to estimate weld
metal fracture toughness, previous research has found that
production welds may have lower fracture toughness than
the procedure qualification record (PQR) test welds due to
differences in the joint geometry, and so use of minimum
Charpy requirements of PQR test welds may be unconservative in some cases [31].
The CVN values were converted into KIc values using
the Charpy/fracture toughness correlation in BS 7910. This
approach utilizes the master curve to estimate the material
toughness. Other factors in this equation are the probability
that the material fracture toughness (Kmat) is less than estimated, the thickness of the material, and a factor (TK), which
describes the scatter in the Charpy versus fracture toughness
correlation. The probability that Kmat is less than estimated
was set to 50% since the Charpy value used in the correlation
is only the minimum specified value and the scatter of the
data is unknown. The thickness was taken as 4″ unless the
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
18
permissible thickness for a certain grade of steel was smaller.
The recommended TK term of +25°C was used to account
for the scatter in the Charpy correlation.
It was found that, except for non-fracture critical Grade
HPS 100W, the base metal Charpy requirements resulted
in lower KIc values than the weld metal requirements. In
order to summarize the data, the minimum base or weld
metal fracture toughness values in AASHTO Temperature
Zone II at the maximum permissible thickness are shown in
Table 7. Grades 36 through 50W have KIc of approximately
46 ksi in for fracture critical welds. Grade 70W has a KIc of
approximately 60 ksi in , and Grade 100W has KIc of approximately 75 ksi in for fracture critical welds. Therefore, use of
50 ksi in for Grade 36 through 50W and 75 ksi in for Grade
100W is probably reasonable. It should be noted that this is
an approximate correlation and the fracture toughness in the
HAZ is unknown.
3.1.3.3 Primary Stresses
The primary stresses (i.e., combined dead and live load
stresses) were assumed to be 75% of the minimum specified
yield strength of the base material. This correlates to the allowable stress level for the operating rating in the 2nd Edition
Manual for Bridge Evaluation [32]. The HL-93 rating factor
for the operating rating is 1.3 times greater than the inventory
rating factor. A reasonable stress level for inventory loading
is 0.55*Fy, since this correlates to the historic allowable stress
design limits. Multiplying 1.3*0.55 approximately results in
the 0.75 factor. This is believed to be a reasonable level for
maximum anticipated loads in a bridge member and that
using 100% of the minimum specified yield strength would
be overly conservative.
3.1.3.4 Residual Stresses
The effects of residual stresses were included in analytical
models for fracture by assuming a uniform tension stress. In
Table 7. Typical fracture toughness values
for bridge welds.
Steel
Non-FC Gr. 36 - HPS 50W
Non-FC Gr. HPS 70W
Non-FC Gr. HPS 100W
FC Gr. 36 - HPS 50W
FC Gr. HPS 70W
FC Gr. HPS 100W
Minimum Base or Weld Metal in Zone II at
Maximum Thickness (ksi in)
NA
>56.51
68.4
45.8
>60.41
>75.11
Note: 1T0 was slightly conservative since minimum Charpy energy was rounded
down to either T27J or T40J.
reality, the actual residual stress profile highly depends on the
type of weld, the welding procedure, etc. Furthermore, the
effects of repairs and starts/stops can drastically affect residual
stress fields. For double-bevel CJP welds, assumed residual
stress profiles typically result in compressive residual stresses in
the root of the weld. This compressive residual stress helps with
resisting fracture for embedded discontinuities, but it results
in different critical crack sizes depending on the location of
the embedded discontinuity with respect to location within the
weld. For new welds, the acceptance criteria would also need
to be different depending on the weld type (e.g., single bevel,
double bevel, electroslag) and the welding procedure (e.g., heat
input). In addition, different criteria would need to be developed for repair welds where the residual stress fields would
also be considerably different. Therefore, a simplified approach
for acceptance criteria for new welds was utilized due to the
additional variables which would need to be included in the
acceptance criteria (i.e., one would require knowledge of heat
input, travel speed, top or bottom of the plate) in order to
account for the different residual stress profiles and the location of the discontinuity in relation to the residual stress profile.
Use of a uniform residual stress profile equal to 100%
of the expected yield strength is recommended by BS 7910
for an initial assumption. BS 7910 includes additional nonuniform residual stress profiles in Annex Q, and API 579-1/
ASME FFS-1 [33] released updated non-uniform residual
stress profiles with the 2016 edition of the code. For both
codes, the residual stress is equal to the assumed material
yield strength at the surface of the weld. For embedded flaws,
the maximum residual stress in the middle two-thirds of the
thickness was evaluated. For BS 7910, this results in a residual
stress of approximately 70% of the assumed material yield
strength. For API 579-1/ASME FFS-1, the residual stress profile depends on the heat input of the weld divided by the weld
thickness, with lower heat inputs or thicker welds resulting
in higher residual stresses. Previous research has found that
typical bridge heat inputs are between 50 kJ/in to 90 kJ/in
[31]. If the heat input was 50 kJ/in, the low heat input
category would be invoked for welds greater than 1.5″ thick.
If the heat input was 90 kJ/in, the low heat input category
would be invoked for welds 3″ thick or greater. For the
low input case, the maximum residual stress in the middle two-thirds of the thickness according to API 579-1/
ASME FFS-1 would be 100% of the assumed material yield
strength. The medium heat input case results in residual
stresses of approximately 60% of the assumed material yield
strength in the middle two-thirds of the thickness.
Two levels of residual stress were evaluated during the parametric study. The first level was residual stresses equal to the
expected yield strength of the base metal (1.1*FyNominal), as it
seems that residual stresses of 100% of the expected material
yield strength are an appropriate assumptions for surface
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
19 flaws. The second level was equal to two-thirds (66%) of the
expected yield strength of the base metal, as it seems that using
66% of the assumed material yield strength is appropriate for
embedded flaws in most cases.
BS 7910 includes an additional equation which accounts
for global relief of residual stresses due to primary loading
of the structure. This equation was utilized in the evaluation
which resulted in a reduction of the actual residual stresses
at the surface from 100% of the expected yield strength to
approximately 75% of the expected material yield strength
for equal thickness butt welds and 60% of the expected yield
strength for thickness transition butt welds.
3.1.3.5 Summary of Parametric Study
Using the inputs of the fracture parametric study, the maximum crack height and length that would resist fracture was
determined for various a/c ratios, yield and tensile strength,
fracture toughness, primary stress, and residual stress using
SignalFFS software [29]. The specific joint types that were
evaluated for fracture are the same as those evaluated for
fatigue shown in Table 6.
The fracture analysis and findings are presented in the following sections, along with the fatigue analysis findings for
each geometry type, including equal thickness butt welds and
thickness transition butt welds. The results of T and corner
joints are presented in Appendix A.
3.1.4 Critical Flaw Sizes for Equal
Thickness CJP Butt Welds
For the analysis of CJP butt welds, the embedded and near
surface flaws were assumed to be centered in a 12″ wide plate.
As expected, increasing the stress range resulted in a large
decrease in the maximum flaw size that would not result
in fatigue crack growth. Likewise, decreasing the fracture
toughness of the material resulted in a large decrease in the
maximum flaw size that can resist fracture. The maximum
flaw size was similar for embedded cracks at the ¼ point and
the midpoint through the thickness of the plate except for
thin plates (i.e., 0.5″ thick) at a 4 ksi stress range or for fracture toughness of 100 ksi in . In these cases, a slight increase
in maximum flaw size is seen for cracks at the midpoint compared with the ¼ point through the thickness of the plate.
Due to the similarities of the ¼ point and midpoint
embedded crack results, the results from the ¼ point were
used for the critical crack size of embedded flaws. It was
found that increasing the thickness of the plate resulted in an
increase in the critical flaw size for both surface and embedded
flaws for both the fatigue and fracture analyses, but the sensitivity in critical flaw size to changes in plate thickness was
much lower for limiting fatigue crack growth than limiting
fracture. In other words, the critical flaw size was very similar
for all thicknesses when accounting for fatigue crack growth.
The results from the fatigue and fracture parametric
studies were combined in order to determine the target critical flaw size. Obviously, there was a wide spectrum of target
critical flaw sizes depending on the inputs. Plots were created
for both 100% and 66% residual stress assumptions for each
plate thickness and for both surface and embedded cracks.
Due to the wide variation in flaw size due to the underlying
input parameter assumptions, the critical crack size data were
grouped based on selected parameters that were reasonable
for typical bridge welds. Based on the discussion of inputs
noted above, the assumptions given in Table 8 were determined to be most reasonable for target critical flaw sizes of
highway bridge welds.
After completing the review and grouping data, simplified tables for the combined fatigue and fracture analysis
of equal thickness CJP butt welds were developed. The corresponding target critical flaw height and length at various
aspect ratios for the cases in Table 8 are shown in the following tables: Table 9 for surface flaws in equal thickness welds,
and Table 10 for embedded flaws in equal thickness welds.
The results in Table 9 and Table 10 were compared to the
closed fracture mechanics equations for a surface crack and
embedded crack in an infinite plate. Using the equations
provided by Anderson [34], the stress intensity factor for a
circular surface crack in an infinite plate is KI = 1.29s a and
KI = 1.13s a where “a” is the entire crack height for the
surface crack and half the crack height for the embedded
crack. These equations can be easily modified for fatigue
crack growth by exchanging DKth for KI and the cyclic stress
range (Sr) for stress (s), resulting in DKth = 1.29 Sr a and
DKth = 1.13 Sr a for surface and embedded cracks, respectively. Table 11 gives the results of the critical size of circular
flaws in an infinite plate for fatigue loading only using inputs
of 2.5 ksi in for DKth and 4 ksi or 8 ksi for Sr. The 8 ksi stress
Table 8. FFS inputs for target critical flaw sizes.
Residual Stresses (% of
Expected Yield Strength)
Yield Stress
Grade 36–50
Grade 70–100
Grade 36–100
Fracture Toughness
Fatigue Stress
Kc=50 ksi in
4 ksi
Kc=75 ksi in
4 ksi
Kc=75 ksi in
8 ksi
100% for surface flaws, 66% for embedded flaws
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
20
Table 9. Target surface flaw in equal thickness weld.
Yield Stress
Fracture
Toughness
Fatigue Stress
Plate Thickness
0.5"
1"
2"–4"
Grade 36–50
Grade 70–100
Grade 36–100
Kc=50 ksi in
Kc=75 ksi in
Kc=75 ksi in
4 ksi
Flaw
Flaw
x
Height
Length
0.059" x 1.076"
0.063" x 0.736"
0.149" x 0.199"
0.063" x 1.044"
0.136" x 0.339"
0.155" x 0.207"
0.064" x 1.059"
0.137" x 0.343"
0.157" x 0.209"
4 ksi
Flaw
Height
8 ksi
Flaw
x
Length
0.037" x 1.061"
0.065" x 0.218"
0.095" x 0.127"
Flaw
Height
x
Flaw
Length
0.024" x 1.082"
0.032" x 0.187"
0.058" x 0.115"
0.063" x 0.084"
Table 10. Target embedded flaw in equal thickness weld.
Yield Stress
Fracture
Toughness
Fatigue Stress
Plate Thickness
Grade 36–50
Grade 70–100
Grade 36–100
Kc=50 ksi in
Kc=75 ksi in
Kc=75 ksi in
4 ksi
Flaw
Height
x
4 ksi
8 ksi
Flaw
Flaw
Flaw
x
x
Length Height
Length
0.061" x 1.020"
0.097" x 1.023"
0.063" x 0.571"
0.125" x 0.314"
0.093" x 0.170"
0.200" x 0.200"
0.143" x 0.143"
0.062" x 1.041"
0.107" x 1.066"
0.063" x 0.963"
0.125" x 0.430"
0.095" x 0.174"
0.251" x 0.251"
0.152" x 0.152"
0.110" x 1.004"
0.063" x 1.046"
0.125" x 0.462"
0.096" x 0.174"
0.262" x 0.262"
0.153" x 0.153"
Flaw
Flaw
Length Height
0.5"
0.141" x 1.010"
0.221" x 0.276"
1"
0.181" x 1.066"
0.254" x 0.462"
2"–4"
0.196" x 1.031"
0.253" x 0.550"
range results given in Table 11 matched the target surface
and embedded flaws for 4″ thick plates given in Table 9 and
Table 10. The 4 ksi stress range results given in Table 11 are
larger than the target surface and embedded flaws for 4″ thick
plates given in Table 9 and Table 10 since these cases were
controlled by fracture rather than fatigue crack growth.
It quickly became apparent that the critical crack sizes were
quite small in some cases. To verify if the approach was yielding reasonable estimates, the simplified results were compared
to the flaw size acceptance criteria in ISO 19285:2017 [17].
This PAUT inspection document includes acceptance criteria
to be determined by either measuring the height and length
of the flaw or by measuring the length and the maximum
amplitude. There are three levels of quality specified in this
document, but Level 2 would typically be used for bridge
welds. Table 12 includes the maximum flaw height and length
for surface and embedded flaws over various plate thicknesses.
For thin plates (i.e., 0.5″), the critical flaw height tends to be
similar to those calculated during this project, but the critical
Table 12. ISO 19285:2017 Level 2 acceptance
criteria.
Plate
Thickness
0.5"
Table 11. Circular flaw in an infinite plate
with fatigue loading only.
Defect Type
Surface
Embedded
4 ksi Stress Range
0.234" x 0.469"
0.612" x 0.612"
8 ksi Stress Range
0.058" x 0.117"
0.153" x 0.153"
1"
2"–3"
4"
Surface Flaws
Flaw
Flaw
x
Height
Length
0.039" x
0.079" x 0.5"
0.039" x
0.079" x 1"
0.079" x
0.118" x 1.969"
0.118" x
0.157" x 2.362"
Copyright National Academy of Sciences. All rights reserved.
Embedded Flaws
Flaw
Flaw
x
Height
Length
0.039" x
0.079" x 0.5"
0.039" x
0.157" x 1"
0.079" x
0.197" x 1.969"
0.118" x
0.236" x 2.362"
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
21 flaw length tends to be a bit longer for the ISO acceptance criteria. As the plates become thicker, the ISO acceptance criteria
allows larger flaws than were calculated during this project.
Overall, the ISO acceptance criteria compares reasonably well
with the critical flaw sizes computed in this project.
ASME Code Case 2235-13 [35] includes an acceptance
criteria based on measurements of flaw height and length.
The acceptance criteria vary by aspect ratio (a/c) and flaw
height (a/t) for thicknesses 1″ and greater. Therefore, at least
three different aspect ratios were evaluated to compare the
range of acceptable flaw sizes. It was found that the results
developed in this study are typically conservative compared
to the sizes in the ASME Code Case for the high-strength
steels or high stress ranges. For thin plates (i.e., 0.5″), the
critical flaw height tends to be within 1⁄16″ for approximately
similar length flaws, but for thick plates (i.e., 3″) the difference between the critical crack sizes developed in this project
and the ASME limits increases and reaches values of approximately 5⁄8″ in some cases. For the 4 ksi stress range cases, the
critical crack sizes compare quite well, with the ASME Code
Case allowing slightly smaller crack sizes for thin plates and
slightly larger crack sizes for thick plates. For the 8 ksi stress
range case, the critical crack size developed in this project is
typically much smaller than the ASME limits. This is due
to the limit placed on fatigue crack growth in this project,
which was not a consideration during the development of
the ASME Code Case 2235 limits [36].
3.1.5 Critical Flaw Sizes for Thickness
Transition CJP Butt Welds
Butt welds with a transition in thickness have a stress concentration at the start of the transition. This stress concentration may decrease the fatigue and fracture resistance of
cracks, especially if they are located near the surface of the
plate. Typical girder flange welds utilize single-sided (offset)
transitions since the web plate remains at a consistent height,
as shown in Figure 4. The butt weld is typically located in
the region with the greatest stress concentration, at the point
where the thin plate begins the thickness transition. The gen-
Thin Plate
Butt Weld
Thick
Figure 4. Typical thickness transition butt weld on
bridge girder flange plate.
eral location of the butt weld and HAZ is visually apparent
in Figure 4. In order to account for this effect, the magnitude
of the stress concentration needs to be determined for various thickness transitions, as shown in the specimen matrix
in Table 6.
Finite element analysis was used to calculate the magnitude of the stress concentration using the commercial
solver ABAQUS. All of the thickness transitions used a 1 to
2.5 slope, which is the maximum allowed by AASHTO
for Category B butt welds. The thickness transitions with
a minimum plate thickness of 0.5″ were modeled as web
plates, while the thickness transitions with a minimum plate
thickness equal to or greater than 1″ were modeled as flange
plates. For the flange plate models, the flanges were assumed
to be 18″ wide while the web was 0.5″ thick by 36″ deep.
For the web plate models, various flange sizes were used to
capture different levels of stiffness.
In order to capture the restraint provided by the web, the
finite element analysis was performed using a 3D model. Typical transition butt weld geometries have a smooth radius as
shown in Figure 4. Therefore, to better represent the actual
conditions, a 2″ radius was used at the change in slope. The
model was loaded with a unit traction in the axial direction
on the thin plate side while the thick plate side was restrained
from movement in the axial direction.
Figure 5 displays the typical profile for stresses in the longitudinal direction of the girder. This figure shows a view
cut through the flange at the point of maximum axial stress.
As expected, there is a concentration of axial stresses at the
radius of the transition to the thin plate. The lowest axial
stresses occur where the thickness transition ends at the
thick plate. The peak stress occurs within the radius, but the
stresses in front of the radius within the thin plate are still
increased in the region very close to the radius. Typically,
the butt weld would be located in the thin plate near the
thickness transition as shown in Figure 4. In this configuration, the far bevel face is located at the start of the transition
radius. The upper portions of the weld near the top surface
will therefore experience increased stresses compared to the
rest of the thin plate.
The results from various thickness transitions were compared by plotting the stress at each integration point along
a path extending from the point of maximum stress concentration vertically through the thickness of the flange.
The location along the thickness of the flange was then
normalized in order to compare the various transition
geometries. As expected, the greatest stress concentration
factor (SCF) occurred at the transitions with the greatest
relative change in thickness, such as 1″ to 2″ or 2″ to 4″.
It should be noted that the stress concentration occurred
within the top half of the plate, and decreased quickly away
from the plate surface.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
22
Max: 1.8
Figure 5. Stresses for 1” to 2” transition weld.
In order to estimate the effect of the stress concentration
on the fatigue resistance of transition welds, a polynomial
trendline was fit to the results from the three transitions with
the greatest SCF—1″ to 1.5″, 1″ to 2″, and 2″ to 4″ transitions—to obtain the following result, where t is the normalized depth in the plate, with t = 0 on the transition-side face
and t = 1 on the opposite face:
SCFFlange (t ) = 1.6587t 4 − 4.8158t 3 + 5.0949t 2 − 3.0826t + 1.7921
The SCF estimation was then multiplied by the various nominal stresses shown in Table 6 to obtain a through-thickness
fatigue stress profile for each nominal stress level. The polynomial stress profiles were then entered into Signal FFS to obtain
the maximum crack size that would not initiate fatigue crack
growth for both embedded and surface cracks.
A separate web thickness transition model was developed
for thickness transitions where the thin plate was less than
1″ thick. Rather than the thickness transition occurring only
on one side of the web, the web plate transition occurred on
both sides of the web so that the web plate remained centered
on the flange. The web plate thickness transition used in this
analysis was 0.5″ to 1″ thick, since this was assumed as the
largest anticipated thickness transition for a web plate. Unit
axial traction was placed on the web and flange on the thin
side of the weld.
It was found that the SCF of web thickness transitions had a
dependency on the relative stiffness of the flange plate. Therefore, the flange size was varied as follows: no flange, 1″ × 6″,
1.5″ × 9″, and 4″ × 18″. The location of the greatest SCF in the
web thickness transition was located near the flange, as shown
in Figure 6. This figure shows axial stresses in the model with
the lower limit on the color palette set to the nominal stress.
When there is no flange, the web SCF is small since the average
stress level is equal to the nominal stress. As the flange stiffness increases, the average stress level through the thickness
of the web increases at the point of maximum SCF. Since the
maximum SCF in web thickness transitions is dependent on
the flange stiffness, the SCF was increased slightly compared
to the highest FE results for the estimation to be used in the
Signal FFS fatigue evaluations in order to be conservative for
possible flange/web combinations. The following equation
displays the resulting estimation of the SCF for web transitions, where t is the normalized depth in the plate with t = 0 on
the transition-side face and t = 1 on the opposite face:
SCFWeb (t ) = 0.8t 2 − 0.8t + 1.4
The equation for the SCF estimation was then multiplied
by the nominal stress values given in the specimen matrix in
Table 6 to obtain a through-thickness fatigue stress profile
for each nominal stress level. The polynomial stress profiles
were then entered into Signal FFS to obtain the maximum
crack size that would not initiate fatigue crack growth for both
embedded and surface cracks.
Similar finite element models were used for fracture resistance analysis as the fatigue resistance. Due to greater applied
stresses in the fracture analysis, plasticity effects needed to be
accounted for through the use of nonlinear FE analysis using
stress-strain curves for ASTM A709 bridge steels. The plasticity
effects flatten the SCF curves near the surface of the thickness
transition due to localized yielding and increase the depth of
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
23 Figure 6. Web SCF for 4” ë 18” flange.
the SCF below the surface. When all of the thickness transitions
were plotted, it was discovered that the thickness transitions
with the highest SCF were nearly linear for points between the
maximum SCF and the midpoint of the plate. Since the critical
flaws will be those on the top surface or ¼-point through the
thickness, the model of the stresses in the FFS approach only
needs to accurately represent the stress profile over the top half
of the plate. Therefore, it was determined that a linear stress
approximation could be used rather than a polynomial. The
linear SCF estimation used for the fracture analysis of transitions
was represented by using the following equations where Sm
is the membrane (nominal) stress and Sb is the bending stress:
Sm = 0.75 Fy
Sb = 0.6 Sm
The values of Sm and Sb were then used in the Signal FFS
software for the primary stresses along with the previously
noted inputs for secondary stresses, thicknesses, fracture
toughness, etc., to calculate the limiting crack size to resist
fracture. Using this approach, the thickness transition CJP
butt weld geometries in Table 6 were modeled as flat plates
with a 12″ width and a linearized primary stress profile for
the primary stresses in the Signal FFS software to calculate the
limiting crack size to resist fracture.
The results for the target surface and embedded weld
flaws in thickness transition welds are given in Table 13
and Table 14, respectively. In these tables, the plate thickness correlates to the thickness of the thinner plate. The
FFS results were determined based on the SCF listed above,
which were appropriate for the largest thickness transitions
Table 13. Target surface flaw in thickness transition weld.
Yield Stress
Fracture
Toughness
Fatigue Stress
Plate Thickness
0.5"
1"–2"
3"–4"
Grade 36–50
Grade 70–100
Grade 36–100
Kc=50 ksi in
Kc=75 ksi in
Kc=75 ksi in
4 ksi
Flaw
Flaw
x
Height
Length
0.039" x 1.051"
0.063" x 0.254"
0.102" x 0.136"
0.032" x 1.057"
0.065" x 0.187"
0.081" x 0.107"
0.031" x 1.041"
0.063" x 0.181"
0.080" x 0.106"
4 ksi
Flaw
Flaw
x
Height
Length
0.028" x 1.029"
0.032" x 0.318"
0.073" x 0.097"
8 ksi
Flaw
Flaw
x
Height
Length
0.013" x 1.010"
0.016" x 0.111"
0.032" x 0.046"
0.024" x 1.057"
0.032" x 0.176"
0.048" x 0.137"
0.061" x 0.082"
0.008" x 1.005"
0.016" x 0.046"
0.019" x 0.034"
0.020" x 0.027"
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
24
Table 14. Target embedded flaw in thickness transition weld.
Yield Stress
Fracture
Toughness
Fatigue Stress
Plate Thickness
0.5"
1"
2"–4"
Grade 36–50
Grade 70–100
Grade 36–100
Kc=50 ksi in
Kc=75 ksi in
Kc=75 ksi in
4 ksi
Flaw
Height
Flaw
x
Length
0.099" x 1.041"
0.126" x 0.349"
0.205" x 0.205"
0.110" x 1.003"
0.125" x 0.448"
0.245" x 0.245"
0.114" x 1.037"
0.125" x 0.521"
0.264" x 0.264"
evaluated. Therefore, these results should be valid for typical thickness transitions, regardless of the thickness of the
thicker plate.
Comparing the results for thickness transition CJP butt
welds to the equal thickness CJP butt welds shown in Table 9
and Table 10, a larger reduction in critical crack sizes was seen
for surface flaws in transition welds than for embedded flaws
in transition welds. The reduction is also much greater for
plates with a thickness 1″ or greater since these plates were
assumed to be flange plates with offset thickness transitions
rather than web plates with centered thickness transitions for
0.5″ plate thickness.
The thickness transition CJP weld results were compared
to ISO 19285 and ASME Code Case 2235. Neither of these
documents has separate acceptance criteria for thickness
transition welds, so accounting for the SCF decreased the
critical crack sizes calculated for thickness transition CJP
welds in this project compared with those allowed in ISO
19285 and ASME Code Case 2235. In general, the target
flaw sizes developed during this project are similar to those
allowed in the other acceptance criteria for 0.5″ and 1″ plate
thicknesses with a 4 ksi stress range. For the other cases,
the target critical crack sizes developed during this project are smaller than those allowed for the other acceptance
criteria, especially for an 8 ksi stress range. Along with not
accounting for the stress concentration of thickness transition welds, the fatigue crack growth failure mechanism was
not considered in the development of ASME Code Case 2235
[36]. It is not known whether either of these factors was
considered during the development of ISO 19285, as no literature was found that documents the development of that
acceptance criteria. The results of T and corner joints are
presented in Appendix A.
4 ksi
Flaw
Flaw
x
Height
Length
0.077" x 1.031"
0.095" x 0.278"
0.127" x 0.196"
0.172" x 0.172"
0.075" x 1.003"
0.095" x 0.250"
0.125" x 0.192"
0.175" x 0.175"
8 ksi
Flaw
Height
x
Flaw
Length
0.038" x 1.097"
0.066" x 0.101"
0.094" x 0.094"
0.035" x 1.004"
0.063" x 0.089"
0.084" x 0.084"
3.2 Round Robin Results
3.2.1 Introduction
A round robin testing program was performed in order
to gain insight into the capabilities of the current technicians in the steel bridge industry and to identify best
practices for improved flaw detection and flaw characterization. The round robin testing program was used to
determine the minimum flaw size that could be reliably
detected with enhanced ultrasonic methods and how the
advanced methods compare with the historical conventional
UT method. Scanning of a set of specimens containing weld
flaws was performed by five PAUT technicians, two TOFD
technicians, and five conventional UT technicians. Data was
only received by four of the five PAUT technicians who participated. Conventional UT and PAUT technicians were qualified according to the requirements in AWS D1.5 as ASNT
Level II for UT and PAUT, respectively. There was tremendous difficulty identifying TOFD technicians to participate
in the round robin testing program due to lack of availability
or lack of equipment. While eleven TOFD technicians were
contacted and five of these technicians initially agreed to take
part, only two technicians actually committed to testing of
the plates. It is believed that this is directly related to the lack
of TOFD technicians in the bridge and building fabrication
industry. The directions and weld bevel drawings supplied
to conventional UT, PAUT, and TOFD technicians who participated in the round robin testing program are provided in
Appendix H.
PAUT technicians were requested to scan the plates in
accordance to the requirements in Annex K and submit the
scan plan details to the research team. Each technician provided details on the number of line scans; location of line
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
25 scans (i.e., face and side of weld scanned and corresponding
index offset); incidence angle range; angular sweep increment;
calibration/TCG block details; equipment and transducer
make, model, and settings; along with any other information included in AWS D1.5 Table K.2. Rather than providing
the same scan plan to the technicians, having the technicians
develop their own scan plan allowed for documentation of the
variations in possible scan results for the same plate scanned
within the requirements of Annex K. For instance, one technician scanned the thickness transition plates with six line
scans, one technician used five line scans, one technician used
four line scans, and one technician used only two line scans.
If the research team had provided a scan plan to the technicians, the scatter of results would be artificially limited since
they would be provided with additional requirements beyond
what is included in Annex K. Also, by providing a scan plan,
the Research Team would have artificially influenced the flaw
rejection results for PAUT by having the technicians either use
an index offset which would maximize or possibly minimize
the amplitude response of the known flaws.
Table 15 provides the details of the flaws included in the
round robin testing program. There was a total of 19 flaws
implanted within the 11 plates that were circulated in the
round robin. Some plates had multiple flaws while others did
not have any flaws. In addition to the testing by the PAUT,
TOFD, and UT technicians, the plates were tested with
Table 15. Round robin flaw
details.
Flaw
ID
Flaw
Type
Height
(in)
Length
(in)
1
2
3
LOF
LOF
LOF
Toe
Crack
Crack
Crack
Crack
LOF
LOF
LOF
LOF
Porosity
Slag
Slag
Slag
Porosity
Slag
Slag
Slag
0.03
0.22
0.20
0.06
0.40
0.40
0.02
0.04
0.43
0.40
0.17
0.45
0.20
0.43
0.23
0.09
0.32
0.16
0.10
0.13
0.24
0.06
0.17
0.80
0.74
0.40
1.00
0.60
0.80
1.00
3.31
0.37
0.18
0.90
3.27
0.49
0.03
3.61
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
digital radiography and the full matrix capture/total focusing method (FMC/TFM) PAUT to aid in determining the
as-built flaw sizes. The digital RT and FMC/TFM results are
also shown in the following sections. It was determined that
digital RT gave the best estimate of the true flaw location and
length for volumetric flaws while FMC/TFM gave the best
estimate of the true flaw height. The digital RT images of each
plate are shown in Appendix D. It should be noted that not all
flaws could be distinguished on the digital RT scans, especially
lack of fusion (LOF) flaws, which do not produce a density
change in the plan view. Testing was also performed using traditional film RT, but it was found that the contrast of the flaws
was poor after digitizing the film. There were also many film
artifacts and scratches. For these reasons, the digital RT images
were exclusively used for flaw location and length sizing.
The FMC/TFM testing was performed blind so as not to
skew the results by providing the intended flaw height. Two
flaws were not detected by the FMC/TFM technician, and
so flaw heights were not reported for these flaws. The flaw
sizes shown in Table 15 are the sizes based on the flaw height
from FMC/TFM and flaw length from digital RT, where this
information was available. If the flaw height or length was
not available based on the FMC/TFM or digital RT results,
the flaw sizes provided by FlawTech for planar flaws or the
intended flaw sizes for implanted volumetric flaws were used.
A few of the implanted volumetric flaws also had unintended
peripheral flaws which were noted on the digital RT scans.
The peripheral flaws were included in the updated flaw location and length measurements. For these flaws, the accuracy
in reported flaw location by the round robin technicians was
considered both with and without the peripheral flaws in
order to capture all possible hits, since some technicians may
have included the peripheral flaws while others may not.
The shear wave velocity of the round robin plates was
measured and compared to a standard AISI 1018 calibration
block to verify whether the plates would be considered acoustically equivalent to typical calibration standards. The shear
wave velocity of all of the round robin plates was found to be
within 0.5% of the AISI 1018 calibration block, which had
a shear wave velocity of 0.1275 in/µs. As will be explained
further, this is within the ±1% tolerance that was determined
to result in negligible error in the amplitude and incidence
angle. The anisotropic ratio (i.e., ratio of shear wave velocity
in the rolled and transverse to rolled direction) was 0.5% and
lower. This is also within the ±1% tolerance that was noted
to result in acoustic anisotropic behavior and beam splitting.
The highlights of the results of the round robin testing
program are briefly summarized in the following sections,
along with a summary of the scatter in height and length
measurements given in Appendix B, and the scatter in
reported amplitude given in Appendix C.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
26
3.2.2 Flaw Detection and Location
The accuracy in reported flaw location was very poor for
many of the PAUT and conventional UT technicians. Thus,
a “hit” (simply defined as the technician noting that they
detected an indication which matched a known flaw) was
originally defined as any reported indication where any part
of the reported indication was within 1 inch along the longitudinal weld axis of any part of a known flaw in the plate. In
other words, the reported indication did not have to line up
with a known flaw or even overlap with a known flaw, as long
as the gap between the extents of the reported indication and
known flaw was less than 1 inch. Clearly this is a very liberal
criterion for counting a reported flaw as a hit, but all technicians, including PAUT, TOFD, and conventional UT, seemed
to struggle with accurately locating flaws. For PAUT and
TOFD technicians, this may be due to inaccuracy in encoder
calibration or incorrectly using the encoder, while for conventional UT technicians it may be due to physical measurement
error. No consideration was made for correctly measuring the
through-thickness location since the flaw depth was not consistently reported (i.e., top of flaw or maximum amplitude).
It was determined not to be too stringent with developing a
criterion for determining whether a reported flaw correlated
to the intended flaw in order to have adequate data for flaw
height and length sizing accuracy, along with variations in
reported amplitude. API RP 2X: Recommended Practice for
Ultrasonic and Magnetic Examination of Offshore Structural
Fabrication and Guidelines for Qualification of Technicians
[37] includes formulas for scoring the performance of ultrasonic technicians when performing a qualification examination. To be correctly located, this document recommends
that the centerline of the reported indication be within the
boundary of the actual indication or within ½ inch of the
actual centerline of the indication (whichever is greater).
This requirement is obviously stricter than the requirement
developed during the initial review of the round robin data
where flaws did not need to overlap at all as long as the gap
between any portion of the reported flaw and actual flaw
was less than or equal to 1 inch.
The hit/miss data are reported using both criteria in
Tables 16–18, with hits labeled as “1” and misses labeled
as “0.” The criteria when the gap between the actual and
reported flaw is less than 1 inch is shown in Table 17, and the
API RP 2X flaw location criteria is shown in Table 18. Since
these tables highlight accuracy of flaw detection, the data is
only shown for the PAUT results where all relevant indications were asked to be reported regardless of the amplitude.
The average hit rate for all of the flaws—planar flaws
(Flaws 1–11) and volumetric flaws (Flaws 12–19)—is
shown in Table 16. The overall hit rate for PAUT, conventional UT, and TOFD was quite similar when using the
1-inch gap criteria for flaw location. For planar flaws, PAUT
had the highest hit rate and conventional UT had the lowest hit rate. For volumetric flaws, TOFD had the highest hit
rate, and PAUT had the lowest hit rate. Under the more stringent criteria included in API RP 2X, the methods that utilize
encoded line scanning had a larger drop in overall hit rate
than the manual, raster scanned conventional UT method.
For planar flaws, TOFD had the highest hit rate, and PAUT
had the lowest hit rate. For volumetric flaws, conventional UT
had the highest hit rate, and TOFD had the lowest hit rate.
The detection rate results for individual flaws were found to
be highly variable. A small slag inclusion (Flaw 18: 0.06″ high
× 0.03″ long) was not detected by any PAUT or conventional
UT technician, but was detected by both TOFD technicians
using the 1-inch gap criteria. It was noted in the inspection
report for one of the TOFD technicians that this flaw had a low
signal response. Another slag inclusion (Flaw 15: 0.10″ high ×
0.90″ long) was detected by every PAUT and conventional UT
inspection, but was missed by one of the TOFD technicians.
3.2.3 Flaw Rejection and Flaw Type
Characterization
Both flaw location criteria (i.e., the 1-inch gap criteria developed by the authors and the API RP 2X criteria) were evaluated
to compare the rejection rate for PAUT Annex K and conventional UT. Flaw classification and rejection rates are shown
in Tables 19–22.
The criteria when the gap between the actual and reported
flaw was 1 inch or less is shown in Table 21, while the API RP
2X flaw location criteria is shown in Table 22, with rejection
labeled as “1” and acceptance labeled as “0” in both tables.
Table 16. Average hit/miss rate.
Location
Criteria
1-Inch Gap
API RP 2X
Method
PAUT
Conventional UT
TOFD
PAUT
Conventional UT
TOFD
Planar Flaws
Only
0.97
0.91
0.95
0.65
0.75
0.91
Volumetric Flaws
Only
0.68
0.83
0.94
0.64
0.83
0.63
All
Flaws
0.86
0.87
0.95
0.64
0.78
0.79
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
27 Table 17. Hit (“1”)/miss (“0”) comparison for 1-inch gap criteria.
PAUT1
Specimen Details
Conventional UT
TOFD
Flaw
ID
Flaw Type
PAUT1
PAUT2
PAUT3
PAUT4
PAUT
Avg
UT1
UT2
UT3
UT4
UT5
UT
Avg
TOFD1
TOFD2
TOFD
Avg
1
LOF
1
1
1
1
1
1
1
1
1
1
1
1
1
1
2
LOF
1
1
1
1
1
1
1
1
1
1
1
1
3
LOF
1
1
1
1
1
1
1
1
1
1
1
1
4
Toe Crack
0
1
0.75
0
0
1
1
1
0.6
1
0
0.5
5
Crack
1
1
1
1
1
1
1
1
1
1
1
1
6
Crack
1
1
1
1
1
1
1
1
1
1
1
1
7
Crack
1
1
1
1
1
1
1
1
1
1
1
1
8
LOF
1
1
1
1
1
1
0
1
1
1
0.8
1
1
1
9
LOF
1
1
1
1
1
1
1
1
0
1
0.8
1
1
1
10
LOF
1
1
1
1
1
1
0
1
1
1
0.8
1
1
1
11
LOF
1
1
1
1
1
1
1
1
1
1
1
1
1
1
12
Porosity
0
0
1
1
0.5
0
1
1
1
1
0.8
1
1
1
13
Slag
0
1
0.5
1
1
1
1
1
1
1
1
1
14
Slag
1
1
1
1
1
1
1
1
1
1
1
1
15
Slag
1
1
1
1
1
1
1
1
1
0
1
0.5
16
Porosity
1
1
1
1
1
0
1
1
0.8
1
1
1
17
Slag
1
1
1
1
1
1
1
1
1
1
1
1
18
Slag
0
0
0
0
0
0
0
0
0
0
0
1
1
1
19
Slag
1
1
1
1
1
1
1
1
1
1
1
1
1
1
Total Hits
15
7
8
18
48
16
15
17
17
18
83
18
18
36
Hit Rate Avg.
0.79
0.78
0.89
0.95
0.86
0.84
0.79
0.89
0.89
0.95
0.87
0.95
0.95
0.95
1Cells
1
1
blacked out were not tested by technician
Flaw rejection from TOFD inspections cannot be compared
to PAUT or conventional UT inspections since there is no
acceptance criteria for this NDT technique in AWS D1.5. For
the flaw to be considered rejected, it had to be located correctly
as well as meeting any other criteria for rejection included in
AWS D1.5. All flaws except Flaw 18, which is the very small
slag inclusion, were considered rejectable by at least one conventional UT technician when using the liberal flaw detection
criteria of the gap being 1 inch or less between the actual and
reported flaw.
The average rejection rate for PAUT and conventional UT
is shown in Table 19 for planar flaws, volumetric flaws, and all
flaws. The rejection rates were lower for PAUT per Annex K
than for conventional UT, especially when considering the
API RP 2X location criteria. The rejection rate for PAUT
and conventional UT are much closer for planar flaws when
using the 1-inch gap criteria while PAUT has a much lower
rejection rate than conventional UT for volumetric flaws. It
was found that while PAUT has a lower rejection rate for the
small planar flaws such as Flaw 1 and 4, all of the planar flaws
0.2″ high by 0.4″ long or larger were rejectable per Annex K
for the 1-inch gap criteria while some of these flaws were not
detected by conventional UT. Due to inaccuracy of locating
these large planar flaws with PAUT, many of these large planar
flaws would not have been rejected by PAUT according to the
flaw location criteria in API RP 2X.
It is important to recognize that the PAUT and conventional UT rejection rates are for the same set of samples and
is providing a direct comparison of the two methods. This
is unlike some previous studies where the rejection rates of
conventional UT and PAUT Annex K were computed using
different sample sets, which could skew the data. It should be
noted that the rejection rates include Flaw 18, even though
this flaw was not rejectable according to any conventional UT
or PAUT technicians. Therefore, this flaw has equally lowered the rejection rate for PAUT and conventional UT, but
the rejection rates of critical flaws would be slightly increased.
Two indications (Flaw 1 and Flaw 14) were rejected by
PAUT technicians, per Annex K, due to being characterized as
cracks rather than due to amplitude and length. The current
AWS D1.5 Annex K10.2 states that “indications characterized
as cracks shall be considered unacceptable regardless of length
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
28
Table 18. Hit (“1”)/miss (“0”) comparison for API RP 2X criteria.
PAUT1
Specimen Details
Conventional UT
TOFD
Flaw
ID
Flaw Type
PAUT1
PAUT2
PAUT3
PAUT4
PAUT
Avg
UT1
UT2
UT3
UT4
UT5
UT
Avg
TOFD1
TOFD2
TOFD
Avg
1
LOF
1
0
1
1
0.75
0
1
0
1
1
0.6
0
1
0.5
2
LOF
1
1
1
1
1
1
1
1
1
1
1
1
3
LOF
1
1
1
1
1
1
1
1
1
1
1
1
4
Toe Crack
0
1
0.5
0
0
0
0
0
0
1
0
0.5
5
Crack
1
0
0.5
1
1
1
1
1
1
1
1
1
6
Crack
1
0
0.5
1
1
1
1
1
1
1
1
1
7
Crack
1
0
0.5
1
1
1
0
1
0.8
1
1
1
8
LOF
1
1
0
1
0.75
1
0
1
1
1
0.8
1
1
1
9
LOF
1
0
0
1
0.5
0
1
1
0
0
0.4
1
1
1
10
LOF
1
0
1
1
0.75
1
0
1
1
1
0.8
1
1
1
11
LOF
1
0
1
0
0.5
1
0
1
1
1
0.8
1
1
1
12
Porosity
0
0
1
1
0.5
0
1
1
1
1
0.8
0
1
0.5
13
Slag
0
1
0.5
1
1
1
1
1
1
0
1
0.5
14
Slag
1
0
0.5
1
1
1
1
1
1
0
1
0.5
15
Slag
1
1
1
1
1
1
1
1
1
0
1
0.5
16
Porosity
1
1
1
1
1
0
1
1
0.8
1
1
1
17
Slag
1
1
1
1
1
1
1
1
1
1
0
0.5
18
Slag
0
0
0
0
0
0
0
0
0
0
0
0
1
0.5
19
Slag
1
1
1
1
1
1
1
1
1
1
1
1
1
1
36
14
14
15
15
16
74
13
17
30
0.64
0.74
0.74
0.79
0.79
0.84
0.78
0.68
0.89
0.79
0
1
Total Hits
15
2
6
13
Hit Rate Avg.
0.79
0.22
0.67
0.68
1Cells
blacked out were not tested by technician
or amplitude.” The rejection rate of PAUT Annex K without
invoking Clause K10.2 (i.e., these flaws would instead be
acceptable due to low amplitude) is included in Table 19, since
PAUT flaw characterization is not always accurate and many
PAUT technicians may not be comfortable with or capable of
characterizing discontinuities as cracks. In fact, both of these
indications were mischaracterized as cracks. Flaw 1 is a small
lack of fusion indication which was acceptable per all other
PAUT inspections but was rejectable to most conventional
UT inspections. Flaw 14 is a slag indication and was accept-
able per the other PAUT inspection but was rejectable to all
conventional UT inspections.
While the rejection rate for PAUT and conventional UT
was similar for many individual flaws, some individual flaws
had much lower rejection rates for PAUT than for conventional UT. There were two reasons for these differences: the
flaw classification (Class A–D) based on amplitude for PAUT
was often lower than conventional UT, and the reported
location of the flaw was often more inaccurate for PAUT as
compared with conventional UT.
Table 19. Average rejection rate.
Location
Criteria
1-Inch Gap
API RP 2X
Method
PAUT
PAUT w/o K10.2
Conventional UT
PAUT
PAUT w/o K10.2
Conventional UT
Planar Flaws
Only
0.79
0.76
0.87
0.53
0.50
0.75
Volumetric Flaws
Only
0.50
0.45
0.83
0.50
0.45
0.83
All
Flaws
0.68
0.64
0.85
0.52
0.48
0.78
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
29 Table 20. Average flaw classification for Flaw 1 and
Flaw 12–15.
Method
PAUT
Conventional UT
No Reported
Flaw
21%
4%
Class D
Class C
Class B
Class A
21%
4%
29%
4%
21%
0%
7%
88%
line scanning where the reported amplitude is not maximized
compared with the manual, raster scanned conventional UT
approach. Another very important reason for these variations is
due to differences in the amplitude limits between the flaw classifications for the PAUT code and the conventional UT code.
In other words, the Class A–D limits for the PAUT code may not
be equivalent to the Class A–D limits of the conventional UT
code, even if the amplitude was maximized for each method.
Another set of flaws, Flaws 5–7 and Flaw 11, shows large
decreases in the rejection rate for PAUT when applying
the more stringent API RP 2X flaw location criteria. This can
be seen when comparing the results for these individual flaws
For instance, flaws with lower classifications according
to PAUT as compared to conventional UT include Flaw 1
and Flaws 12–15. These include a small LOF flaw, a group
of porosity flaws, and three slag inclusions. In order to illustrate this difference, the average reported flaw classification
for these five flaws are shown in Table 20. PAUT technicians
had a much higher scatter in how these five flaws were classified. As shown in the table, the classification ranged from
Class B, C, or D to not reporting the flaw at all. However, these
flaws were typically reported as Class A for conventional UT.
It is thought that this difference is mainly due to the fact that
PAUT scanning in accordance with Annex K is performed by
Table 21. Rejection rate (reject “1”/accept “0”) for 1-inch gap criteria.
PAUT Annex K1
Specimen Details
Conventional UT
Flaw
ID
Flaw Type
PAUT1
PAUT2
PAUT3
PAUT4
PAUT
Avg
UT1
UT2
UT3
UT4
UT5
UT
Avg
1
LOF
0
0
0
12
0.25
0
1
1
1
1
0.8
2
LOF
1
1
1
1
1
1
1
1
1
3
LOF
1
1
1
1
1
1
1
1
1
4
Toe Crack
0
0
0
0
0
1
0
1
0.4
5
Crack
1
1
1
1
1
1
1
1
1
6
Crack
1
1
1
1
1
1
1
1
1
7
Crack
1
1
1
1
1
1
1
1
1
8
LOF
1
1
1
1
1
1
0
1
1
1
0.8
9
LOF
1
1
1
1
1
1
1
1
0
1
0.8
10
LOF
1
1
1
1
1
1
0
1
1
1
0.8
11
LOF
1
1
1
1
1
1
1
1
1
1
1
12
Porosity
0
0
0
1
0.25
0
1
1
1
1
0.8
13
Slag
0
1
0.5
1
1
1
1
1
1
14
Slag
12
0
0.5
1
1
1
1
1
1
15
Slag
0
0
0
1
1
1
1
1
1
16
Porosity
1
1
1
1
1
0
1
1
0.8
17
Slag
1
1
1
1
1
1
1
1
1
18
Slag
0
0
0
0
0
0
0
0
0
0
0
19
Slag
1
1
1
1
1
1
1
1
1
1
1
Total Rejected
13
5
5
15
38
15
15
17
16
18
81
Rejection Rate
0.68
0.56
0.56
0.79
0.68
0.79
0.79
0.89
0.84
0.95
0.85
1Cells
0
0
blacked out were not tested by technician
due to crack classification rather than amplitude and length
2Rejected
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
30
Table 22. Rejection rate (reject “1”/accept “0”) for API RP 2X criteria.
PAUT Annex K1
Specimen Details
Conventional UT
Flaw
ID
Flaw Type
PAUT1
PAUT2
PAUT3
PAUT4
PAUT
Avg
UT1
UT2
UT3
UT4
UT5
UT
Avg
1
LOF
0
0
0
12
0.25
0
1
0
1
1
0.6
2
LOF
1
1
1
1
1
1
1
1
1
3
LOF
1
1
1
1
1
1
1
1
1
4
Toe Crack
0
0
0
0
0
0
0
0
0
5
Crack
1
0
0.5
1
1
1
1
1
1
6
Crack
1
0
0.5
1
1
1
1
1
1
7
Crack
1
0
0.5
1
1
1
0
1
0.8
8
LOF
1
1
0
1
0.75
1
0
1
1
1
0.8
9
LOF
1
0
0
1
0.5
0
1
1
0
0
0.4
10
LOF
1
0
1
1
0.75
1
0
1
1
1
0.8
11
LOF
1
0
1
0
0.5
1
0
1
1
1
0.8
12
Porosity
0
0
0
1
0.25
0
1
1
1
1
0.8
13
Slag
0
1
0.5
1
1
1
1
1
1
14
Slag
12
0
0.5
1
1
1
1
1
1
15
Slag
0
0
0
1
1
1
1
1
1
16
Porosity
1
1
1
1
1
0
1
1
0.8
17
Slag
1
1
1
1
1
1
1
1
1
18
Slag
0
0
0
0
0
0
0
0
0
0
0
19
Slag
1
1
1
1
1
1
1
1
1
1
1
1Cells
0
0
Total Rejected
13
2
3
11
29
14
14
15
15
16
74
Rejection Rate
0.68
0.22
0.33
0.58
0.52
0.74
0.74
0.79
0.79
0.84
0.78
blacked out were not tested by technician
due to crack classification rather than amplitude and length
2Rejected
in Table 21 and Table 22 for PAUT and conventional UT. These
flaws include three cracks and one LOF flaw. If one were to
group the results of these flaws, the average rejection rate
for PAUT decreased from 100% when using the more liberal
1-inch gap criteria to 50% when using the API RP 2X criteria,
while the conventional UT results only decreased from 100% to
90% using the same criteria, respectively. Therefore, poor flaw
location of these planar flaws resulted in a large decrease in
the rejection rate for PAUT compared with conventional UT.
Table 23 shows the findings for all the PAUT flaw characterizations during the round robin testing program. Cracks were
only correctly characterized 22% of the time, while many false
calls were incorrectly characterized as cracks. It should also
be noted that planar flaws such as cracks and lack of fusion
were sometimes incorrectly characterized as volumetric discontinuities (11% and 4%, respectively). This is a major concern for developing separate acceptance criteria for planar and
volumetric discontinuities. If a planar flaw is mischaracterized
as volumetric, it would result in an unconservative evaluation.
Table 24 compares the flaw characterization results for PAUT
and TOFD. It was found that the average TOFD technicians did
not classify flaws as well as the average PAUT technician. While
the classification of planar flaws was similar, TOFD more
often incorrectly classified volumetric flaws as being planar.
3.2.4 False Calls
The reported indications that were not within ±1 inch of
the total extents of the actual flaw were initially determined
to be false calls. The digital RT results revealed that some of
the plates included unintended weld flaws, especially near the
edge. This was especially true for the FlawTech plates, which
Table 23. Flaw characterization using PAUT.
Actual
Flaw Type
Crack
Crack
LOF
Porosity
Slag
False Calls
22%
21%
25%
9%
71%
Reported Flaw Type
Planar
Volumetric
(Non-Crack)
44%
11%
71%
4%
25%
50%
64%
27%
0%
29%
Copyright National Academy of Sciences. All rights reserved.
No Type
Reported
22%
4%
0%
0%
0%
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
31 Table 24. Comparison of flaw characterization of PAUT
and TOFD.
Reported Flaw Type
PAUT
TOFD
No Type
No Type
Planar Volumetric
Planar Volumetric
Reported
Reported
85%
6%
9%
86%
10%
4%
Planar
60%
0%
86%
0%
14%
Volumetric 40%
71%
29%
0%
100%
0%
0%
False Calls
Actual
Flaw Type
seemed to have groups of sparse porosity intermittently within
some of the plates, as shown in the digital RT images in
Appendix D. Therefore, indications that overlap with these
unintended weld flaws should not be indicated as false calls.
It was found that the number of false calls which the technicians classified as rejectable was relatively small after removal
of these unintended weld flaws which were apparent on the
digital RT results as shown in Table 25.
3.2.5 API RP 2X Scoring Results
API RP 2X includes formulas for scoring the performance
of ultrasonic technicians during a qualification examination.
These formulas evaluate the detection of indications and
include a penalty for false calls. It includes suggested minimum performance scores for qualification examinations. The
performance scores are defined as follows:
P=
Lc
× 100
La
Formula 1
Lc
Lf
R =    1 −  × 100
 L1  
L1 
Formula 2
Where:
P =percentage of actual reflectors correctly detected and
sized
R = overall rating including penalty for false calls, 0 to 100
Table 25. Rejectable false calls.
Technician
PAUT1
PAUT2
PAUT3
PAUT4
UT1
UT2
UT3
UT4
UT5
TOFD1
TOFD2
Number of
Rejectable False
Calls
1
0
0
3
0
1
0
0
0
2
0
Total Length of
Rejectable False Calls
3.11"
0"
0"
2.76"
0"
0.45"
0"
0"
0"
0.95"
0"
La = length of actual reflector contained in the test plate
Lc =credited length for indications that have been correctly
sized and located. (Credit is given for the lesser of the
reported length or actual length of the flaw.)
L1 =accumulative length of all indications by the technician, right or wrong
Lf =accumulative length of indications above the stated
disregard level where no reflector exists
To be correctly sized, API RP 2X recommends that the
reported dimensions be within a factor of two of true dimensions (i.e., one-half to twice the actual dimension). To be correctly located, it is suggested that the centerline of the reported
indication be within the boundary of the actual indication
or within ½ inch of the actual centerline of the indication
(whichever is greater).
API RP 2X suggests that minimum performance for ultrasonic technicians be a score of 70 or above for Formula 1
and a score of 50 or above for Formula 2. These criteria were
applied to the round robin data to compare the performance
of PAUT and conventional UT technicians. The small slag
inclusion (Flaw 18), which was not detected by any PAUT or
conventional UT technicians, was not included in the analysis
since it is not likely to be critical for any bridge structures.
The data was analyzed two ways: (1) only accounting for
correctly measuring flaw length along with flaw location
and (2) accounting for correctly measuring flaw length and
flaw height along with the flaw location. Due to the presence
of unintended peripheral indications outside of the extent
of the intended indications, technicians were given credit
if the reported location and dimensions were within the
required tolerances for either the main grouping (intended)
of the indication or the total flaw including the unintended
peripheral indications. The length of the total flaw, including
unintended peripheral indications, was used for the actual
length (La). Table 26 displays the results considering detected
discontinuities—even if they were not deemed rejectable—
and disregarding height sizing errors for PAUT and TOFD
technicians. No PAUT or TOFD technicians met the minimum requirements for P and R, but two conventional UT
technicians passed both requirements. The average score was
much lower for the PAUT and TOFD technicians than the
conventional UT technicians. Table 27 displays the results for
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
32
Table 26. API RP 2X results, not including height sizing error.
TOFD1
TOFD2
Avg. TOFD
P
47
65
56
R
39
82
61
PAUT1
PAUT2
PAUT3
PAUT4
Avg. PAUT
PAUT and TOFD when length and height sizing were both
required to be within one-half to twice the actual dimension,
along with being properly located. Due to the large inaccuracies with height sizing and the lack of reported flaw height for
TOFD results, no PAUT or TOFD technicians were close to
passing the minimum performance requirements.
This simple performance test highlights that there is considerable room for improvement to the current PAUT procedures. For instance, if one were to compare the numerical
values to what a student might obtain in a college engineering
course, the grades for PAUT technicians are generally in the
“D” to “F” range at best for Table 26 and at worst when considering Table 27. In the opinion of the authors, this strongly
highlights the need for performance testing of ultrasonic
technicians in the bridge industry in order to test their abilities to detect (and reject) critical weld flaws.
3.2.6 Modifications to Future Scanning
Procedures Based on Round
Robin Results
The initial round robin testing phase proved to be extremely
useful and revealed much about the state of the practice as
related to UT and PAUT. However, based on the results of the
round robin, moving forward with development of acceptance criteria based solely on the physical measurement of
the flaw size was deemed not to be feasible with the present
accuracy and reliability of flaw size measurements using the
current PAUT workforce. Acceptance criteria based on flaw
height and length sizing was still provided as an alternative
method if the PAUT technician develops a written procedure
according to specified requirements and successfully comTable 27. API RP 2X results for PAUT
and TOFD, including height sizing.
TOFD1
TOFD2
Avg. TOFD
P
17
29
23
R
14
37
26
PAUT1
PAUT2
PAUT3
PAUT4
Avg. PAUT
P
29
10
20
33
23
R
24
11
25
26
22
P
68
44
48
64
56
R
56
50
60
50
54
UT1
UT2
UT3
UT4
UT5
Avg. UT
P
67
75
61
79
78
72
R
78
86
57
59
48
66
pletes a performance test on samples of similar material and
with flaws similar to the critical flaw size. This would allow
for other advanced methods such as FMC/TFM (Full Matrix
Capture/Total Focusing Method) PAUT to be used in inspecting bridge welds provided that they can detect and reject critical weld flaws if the shop, technician, engineer of record, etc.,
would request such methods.
The conclusion was also that implementation of TOFD will
be very difficult in the bridge industry. TOFD testing is specialized and not readily available; in addition, it is clear that
there are very few TOFD technicians that perform testing on
structural welds, and it appears that there is not much interest
in testing bridge welds with TOFD. TOFD technicians noted
that testing butt welds where there is a plate thickness transition is also difficult with TOFD due to the special attention
and jigs needed to scan these welds.
The focus of the research, according to the research objectives, was to develop acceptance criteria based on maximum
amplitude and flaw length for PAUT in the AWS D1.5. From
the round robin results, it appeared that PAUT, when performed in accordance with Annex K, shows improved detection of planar flaws as compared with conventional UT since
data collected showed that PAUT did not miss large flaws
when using the encoded line scans and the reported flaw
location only had to be within 1 inch of the actual flaw to be
counted as a hit. However, it seems that additional training
or performance testing is required to improve the accuracy
of encoded scans. The testing also confirmed that the lack of
raster scanning, to maximize the amplitude associated with
a given indication, reduces the rejection rate that results from
using the current version of Annex K compared with using
conventional UT. In other words, flaws that are rejected today
using conventional UT would not be rejected using Annex K,
although there is no fitness-for-service or engineering basis
for allowing this to be the case.
Again, while ideally one would prefer to reject the flaw
based on the size of the flaw and the type of flaw (planar vs.
volumetric), the round robin testing shows that neither of
these can be reliably performed with the current workforce.
To ensure that flaws that would be rejected today using conventional UT are not accepted with PAUT using line scanning alone, additional raster scanning of selected indications
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
33 coupled with line scanning using an encoder is prudent. This
will help ensure that the small flaws that were accepted in
the round robin when using line scanning alone would be
rejected in practice. PAUT has the advantage of producing
ultrasound over a range of angles, so additional raster scanning would not only maximize the amplitude at that angle
but would do so over a wide range of angles. This should also
help improve the rejection rate of volumetric flaws as well as
small planar flaws. Further, use of an encoder would allow for
a permanent record of the initial line scan.
In its most basic form, the proposed approach for inspection is to scan a weld using PAUT using an encoder along a
specific index offset to be set during the development of the
scan plan. If an indication is identified that meets a certain
threshold, the technician would return to that location and
raster scan using the PAUT probe in order to maximize the
length and decibel reading. Then, the decision to accept or
reject the indication would be based on an amplitude and
length table.
The round robin testing showed conventional UT is already
being used to reject very small planar flaws (0.03″ × 0.06″).
Thus, it does not seem that the proposed approach will
result in increased rejection rates for small flaws (when using
PAUT), which would have been acceptable under conventional UT and thereby resulting in an unreasonable increase
in weld repairs. However, as discussed previously, while the
line scanning approach is generally repeatable, the variability
of the manual scanning approach after an indication is found
using PAUT is problematic.
3.3 CIVA Modeling
To better minimize the variability found during the round
robin testing program, simulations were performed using
CIVA-UT [38] to aid in the initial procedure development
through modeling. The modeling incorporated weld flaws
similar to the critical flaw sizes developed during the analytical program. These flaws serve as a “lower-bound” flaw
set from which improved acceptance criteria were developed
to consistently reject these flaws. As long as flaws of this
size or larger are consistently rejectable, the procedure can
be deemed effective at removing critical flaws from service.
Therefore, the acceptance criteria are grounded in fracture
mechanics but will not use flaw height measurement for
evaluation.
Two parametric modeling studies were performed using
CIVA-UT. The first parametric study addressed the differences in reference amplitude possible due to differences in
acoustic properties of various steel bridge base metal grades.
This study involved (1) development of benchmarked CIVA
material models based on experimental test results and
(2) CIVA modeling of various probes and incidence angles
to develop recommendations for AWS D1.5 in order to
limit the amplitude differences between the calibration block
and test object.
The second parametric study using CIVA evaluated the
effects of variations in the amplitude response of weld
flaws deemed critical per the FFS parametric study. This
analysis provides a rational comparison of the amplitude
from the target critical weld flaws to the acceptance criteria amplitude limits. These data were used to develop a
rational fracture-mechanics-based PAUT inspection procedure and acceptance criteria which will detect and reject
critical weld flaws. The parametric models varied the plate
thickness along with the flaw type, size, position, tilt, and
skew of target critical weld flaws in order to compare the
maximum amplitude of the indication response with the
reference amplitude and the associated acceptance criteria
amplitude limits.
The detailed results of the two CIVA parametric studies
are included in Appendix E. The results of the first parametric study were used to develop recommendations for AWS
D1.5 Annex K to account for differences in acoustic attenuation and shear wave velocity, as discussed in Section 3.4 and
3.5. The results of the second parametric study were used to
develop recommendations for AWS D1.5 Annex K for scanning procedure requirements and acceptance criteria which
will result in improved detection and rejection of critical
weld flaws as discussed in Section 3.7 and 3.8.
3.4 Acoustic Attenuation
3.4.1 Experimental Testing
3.4.1.1 Experimental Testing Program
Unexpected but significant observations were made during
pilot experimental testing to evaluate the effects of coatings. After removing the paint coatings and grounding the
steel surface smooth, a more than 4 dB difference in material
attenuation was found between side-drilled holes (SDH)
located at a depth of 0.5″ in six different steel specimens
when using a 5 MHz shear wave probe. Currently, the
AWS D1.5 Bridge Welding Code procedures (for conventional UT and Annex K) explicitly assume that all carbon
bridge steels possess the same attenuation characteristics
and no correction or consideration needs to be taken during the inspection of bridge welds. As a result, the difference
in attenuation found during this sequence of testing directly
led to a controlled experimental evaluation of the ultrasonic attenuation in different base metals typically used in
bridge construction. The objective was to investigate the
effect of different variables—such as ultrasonic frequency,
wave mode, material microstructure, and material acoustic
velocity—on the magnitude of material attenuation.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
34
Nine steel specimens were fabricated and tested using conventional UT and PAUT. Table 28 outlines the samples tested
and their properties. Two specimens—ID 36 and ID 50 from
the initial testing that looked at the effects of bridge paints
on PAUT—were reused during this portion of the study.
These specimens represented the least and most attenuating
specimens previously tested within the field. To fully evaluate
these differences in a more controlled setting, samples were
cut from the girders and brought into the laboratory. ID
36 was an “historical” A36 steel; ID 50 was a modern A709
Gr. 50 steel. An additional seven modern high-performance
steels (HPS) were added to further extend the evaluation. The
addition of these new specimens set out to further evaluate
if there were differences in the ultrasonic attenuation characteristics in different plates. Three of the new specimens were
of the quenched and tempered (QT) variety at the mill, while
four of the new specimens were produced using the thermomechanical control process (TMCP).
A review of previous literature [39–41] has confirmed that
chemical composition, grain size, and microstructure have
all been found to affect the acoustic properties and propagation of sound through material. The chemical composition of each specimen was obtained and found to meet the
requirements of its respective ASTM steel standard. For
each specimen, a metallurgical analysis of the grain size and
microstructure was also performed by an outside consultant. Analysis showed the following: Figure 7 shows the grain
structure perpendicular to the rolled direction for each specimen at the central region of the plate magnified at 100X with
Nital etchant.
• Specimen 36 consisted of a Widmanstätten pattern of
ferrite and pearlite.
• Specimen 50 consisted of ferrite and pearlite.
• Specimen 70 had a general structure of fine acicular ferrite
with small spherical carbides, but also visible were bands of
ferrite and low-carbon martensite and bainite.
• Specimens 101 and 102 consisted of QT martensite.
The TMCP specimens all had a variation in grain structure
near the surface in comparison with the central regions.
• Specimen TMCP 1 had acicular ferrite with elongated
pearlite and long bands of pearlite in the central region.
On the near surface region, a fine acicular ferrite and short
bands of pearlite were seen.
• Specimen TMCP 2 had elongated ferrite with bands of
pearlite and bainite in the central region. On the near surface region, elongated ferrite and short bands of pearlite
and bainite existed.
• Specimen TMCP 3 had a fine acicular ferrite with patches
of pearlite in the central region. On the near surface region,
a more refined structure of fine acicular ferrite and patches
of pearlite were seen.
• Specimen TMCP 2 was further analyzed parallel to the
rolled direction. Parallel to rolled, the central region and
near surface regions both consisted of elongated ferrite
with bands of pearlite and bainite.
Figure 7 shows the grain structure perpendicular to the
rolled direction for each specimen at the central region of the
plate magnified at 100X with Nital etchant.
Grain size measurements were made in accordance with
ASTM E112-13 Standard Test Methods for Determining Average Grain Size. Per ASTM E112, grain size measurements
can be conducted numerous ways, but all methods include
counting the number of grains or number of grain boundaries along a specified line within a known area. A table is
provided in ASTM E112 to rate the grain size from 00 up to
14.0, with 00 representing the largest average grain size and
14.0 the smallest average grain size [42]. Table 29 presents the
grain sizes measured for the group of specimens tested for
this project. It should be noted that in this table, for Specimens 101 and 102, the prior austenite grain size is measured
and presented. In this case, the prior austenite grain size was
that of the steel before quenching and tempering occurred.
Table 28. Steel specimens.
ID
36
50
70
101
102
TMCP 1
TMCP 2(2)
TMCP 3
Steel Properties
A36
A709 Gr50
HPS 70W QT
HPS 100W QT
HPS 100W QT
HPS 70W TMCP
HPS 70W TMCP
HPS 70W TMCP
Fabrication Year
1973(1)
2013
2015
circa 2000s
circa 2000s
2009
2014
2011
Thickness (in)
1.25
1.25
1.50
2.00
1.50
1.57
1.25
2.00
Width (in)
1.87
1.87
1.87
1.87
1.87
1.87
1.87
1.87
(1)
Date the bridge was put into service
Two specimens, one in the rolled direction and one in the transverse to rolled direction, were
fabricated with this steel plate
(2)
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
35 36
50
70 (QT)
101 (QT)
102 (QT)
TMCP 1
TMCP 2
TMCP 3
Figure 7. Microstructure for specimens at 100X (perpendicular to rolled direction).
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
36
3.4.1.2 Summary of Experimental Findings
Table 29. ASTM grain size classifications.
ID(1)
Grain Size
36
50
70 (QT)
101 (QT)
102 (QT)
ASTM 2-1/2
ASTM 7
ASTM 10
ASTM 8
ASTM 8
ASTM 11 (central)/
ASTM 11 (surface)
ASTM 11 (central)/
ASTM 11 (surface)
ASTM 10 (central)/
ASTM 10 (surface)
ASTM 8 (central)/
ASTM 12 (surface)
TMCP 1
TMCP 2
TMCP 2
(parallel to roll)
TMCP 3
(1)
(2)
ASTM Grains per Unit
Area (in2) at 100X(2)
2.83
64.0
512.0
128.0
128.0
1024.0
1024.0
512.0
128.0/
2896.3
Perpendicular to rolled direction unless noted otherwise
ASTM E112, 2013
The variability in ultrasonic inspection of NGI-ESW welds
was then assessed following the evaluation of base metal. Unlike
the consistent microstructure of base metal, welding produces
different zones of varying grain structures. From the electroslag
welding process, the HAZ consists of two grain structure zones.
The portion of the HAZ bordering the base metal is comprised
of fine grains and the inner portion is comprised of coarse
grains. The weld may have an additional two or three zones
itself of coarse columnar and/or equiaxed grains [43]. Specimens were fabricated to facilitate the comparison of attenuation
among base metal, HAZs, and weld metal. The electroslag weld
samples were donated by the Federal Highway Administration
(FHWA) and supplied by two different fabricators. As shown in
Figure 8, the individual grain structure of the NGI-ESW welds
varied considerably for these specimens. Table 30 outlines the
details of the specimens and their material properties.
The specific details of the experimental testing program
which measured the acoustic attenuation of base metal and
NGI-ESW are given in Appendix F. The results of these tests
are summarized following.
During the portion of the research that was intended to
investigate the effects of coatings on scanning, major differences were observed in the attenuation characteristics
in different steel plates. In has been generally assumed by
the industry that the attenuation characteristics of the test
block can be used to represent those of the test component. However, upon closer study and laboratory testing,
it was concluded that this assumption is not appropriate in
all cases. Some steels will have much different attenuation
characteristics than that of the typical test block. Clearly, if
the test piece (and weld) are more attenuating than the test
block, the amount of sound returned from a reflector will
be less than the same reflector in the test block. The result is
obvious: unacceptable flaws will be accepted. The converse
is also true if the test block is more attenuating than the test
component.
One of the issues that is specific to PAUT (and Annex K)
is the fact that the effect of attenuation error becomes
more pronounced at higher transducer frequencies. Currently, Annex K permits the transducer frequencies ranging from 1 MHz to 6 MHz, while D1.5 Clause 6 limits the
frequency to 2.25 MHz for conventional UT. Interestingly,
other specifications openly acknowledge this issue and
require technicians to calibrate using a test block made of
a material with the same acoustic properties as the piece
being inspected unless it is confirmed that the attenuation
characteristics are comparable as previously stated. This
would, of course, include attenuation associated with the
weld metal.
In lieu of requiring such calibrations, the experimental
results suggest that for most cases, these effects in base metal
will generally become small and possibly negligible when
the transducer frequency is less than or equal to 2.25 MHz.
The trade-off is that lower-frequency transducers have less
resolution, but the CIVA and experimental studies suggest
this will not adversely affect the overall reliability.
Figure 8. Variation in NGI-ESW grain structure.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
37 Table 30. NGI-ESW specimen.
Specimen ID
Fabricator
P1
P2
P3
P4
Fabricator A
Fabricator A
Fabricator A
Fabricator B
Base Metal
Side A
Side B
HPS 70W (QT) HPS 70W (QT)
50W
50W
50W
HPS 70W (QT)
HPS 70W
HPS 70W
While limiting the probe frequency may address the
attenuation issues associated with base metal, the attenuation characteristics of the weld itself, in particular attenuation associated with NGI-ESW, was found to be quite high.
In this case, use of a lower frequency, such as 2.25 MHz,
still resulted in a large loss of amplitude from sound passing through the weld metal. In other industries, when the
attenuation of base metal or weld metal is suspected as
being different than the calibration block, a transfer correction is formulated through a pitch-catch scanning procedure to measure the actual attenuation. It is clear that the
material attenuation issues observed must be accounted for
in some way.
3.4.2 AWS Recommendations
for Acoustic Attenuation
Differences in acoustic attenuation for common grades
of base metals were noted during the experimental testing.
The differences in attenuation were significantly greater for
5 MHz probes than 2.25 MHz probes, which resulted in large
amplitude variations depending on the differences between
the calibration block and the test object. CIVA models representing various grades of bridge base metals were developed
and benchmarked to the experimental test results. The CIVA
models were used to evaluate the amplitude differences that
would result from testing different grades of base metals after
performing calibration on a typical calibration block with
AISI 1018 steel base metal.
Based on this analysis, proposed modifications to AWS
D1.5 Annex K were developed in order to improve the calibration requirements to account for the differences in acoustic
attenuation. A marked-up version of Annex K and associated commentary with the recommended changes is given
in Appendix G. These modifications include that a 2.25 MHz
probe should be used for plate thicknesses exceeding 0.5″
unless the attenuation of the calibration block and test object
is similar or the differences are accounted for through the
use of a transfer correction. A transfer correction accounts
for the differences in attenuation and coupling losses due
to surface roughness between the calibration block and the
test object. Transfer correction is referred to in many UT
references [12, 13, 19, 37, and 44] specifically as a method
Fabrication
Year
2015
2015
2015
2013
Thickness (in)
2.0
2.0
2.0
2.0
to account for acoustic differences between the calibration
block and the test object. Two methods may be used to
perform a transfer correction:
1. Fabricate a block with the same reference reflector including type, size, and depth as the calibration block, and note
the differences in the signal response between the calibration block and the test object. The sound path for this
measurement should correspond to the longest sound
path used for the inspection.
2. Perform pitch–catch measurement on the calibration
block and the test object over two sound paths to develop
the relationship between attenuation and sound path that
can be used to correct the amplitude from the calibration
block.
Since the first option involves fabrication of a new calibration block, the second option is often much more suitable
for checking whether any amplitude correction is necessary
between an existing calibration block and the test object. In
this method, two probes are used, with one acting as a transmitter and the other as a receiver. The sound is skipped off
of the backwall of the calibration block and test object in a
single-V and double-V path (i.e., one skip and two skips off
of the backwall), as shown in Figure 9, and the amplitude is
measured at each location with the same transducer settings.
These amplitude measurements are then plotted against the
sound path and lines are drawn through the measurements
corresponding to the calibration block and the test object,
as shown in Figure 10.
The difference in amplitude at the maximum inspection
sound path can be directly obtained from these lines. If the
calibration block line is above the test object line, then the
calibration block has less attenuation than the test object and
additional gain must be added to the inspection of the test
object. If the calibration block line is below the test object
line, then the calibration block has more attenuation than
the test object. While gain may be removed from the inspection of the test object to correct for the difference, in this
case, this could result in removal of too much gain at shorter
sound paths.
The thickness of the calibration block and the test object
does not need to match, but the amplitude measurements must
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
38
Figure 9. Transfer correction probe locations.
be made in the far field in order to ensure that the change
in amplitude is due to attenuation. API RP 2X requires that
the single-V-path-measurement sound path exceeds 4 inches.
If the single-V-path measurement is within this sound path,
it is recommended that additional skip distances be used to
obtain longer sound paths.
While the inspection will be performed with PAUT probes,
use of conventional UT probes with similar frequency and
aperture that will be used for the PAUT inspection will be
adequate to account for the attenuation. It is recommended
that the conventional UT wedge be chosen to produce the
largest incidence angle used in the PAUT inspection, since
this will increase the sound path and is more sensitive to
amplitude differences due to velocity variations. During
experimental testing, 2.25 MHz and 5 MHz conventional UT
and PAUT probes were used for attenuation testing, and the
results of PAUT and conventional UT were very similar for
the same incidence angle, frequency, and aperture.
According to ISO 17640 [19], no correction is required if
the greatest difference in amplitude between the calibration
block and the test object at the longest sound path is 2 dB or
less. This code requires correction for amplitude differences
greater than 2 dB but smaller than or equal to 12 dB. If the
amplitude difference exceeds 12 dB, this code requires that
the reason for this difference shall be considered or the scanning surface reconditioned.
In order to make the transfer correction when TCG is used,
either (1) the entire TCG curve is corrected by modifying
Figure 10. Transfer correction amplitude.
the reference gain, which would offset the entire TCG curve
up/down by the same amount, or (2) separate corrections could
be performed manually to each TCG point for each focal law.
The first method is very simple, but may result in overcorrection
at short sound paths. For instance, shifting the entire amplitude
by the transfer correction at the longest sound path in Figure 10
would result in the test object amplitude being overestimated at
shorter sound paths. This would be conservative, but the opposite is true if the transfer correction resulted in removal amplitude (i.e., too much amplitude would be removed at short sound
paths resulting in lower sensitivity than reference). Therefore, in
order to avoid unconservative corrections, Annex K should not
allow for the removal of amplitude unless consideration is given
for the amplitude difference at shorter sound paths.
The second method of TCG correction for attenuation
differences described above involves manual adjustment to
each TCG point for each focal law. While this would result
in accurate correction over all sound paths, it would be a
time-consuming process that would require input of at least
93 corrections considering 31 focal laws for a 40°–70° incidence angle range with 1° increment and a 3-point TCG.
Therefore, unless automated in some way, this method is
more sensitive to human error and unlikely to be practical
for shop application.
The research team recommends that AWS Annex K be
modified to require verification of the differences in acoustic
properties between the test object and calibration block. If
the amplitude difference exceeds 2 dB at the longest sound
path and the calibration block is less attenuating than the test
object, a correction for this difference is required. Guidance
for four methods of correction has been produced for the
AWS Annex K commentary. They involve the following:
1. Using a lower frequency probe, such as 2.25 MHz.
2. Using a calibration block that better matches the test
object and results in less than a 2 dB amplitude difference
at the longest sound path.
3. Reconditioning the scanning surface of the test object in
order to better match the calibration block.
4. Adding gain equal to the difference between the calibration block and the test object at the longest sound path.
It is noted that this may be overly conservative at shorter
sound paths and it is not recommended to be performed
for differences greater than 12 dB.
It is very likely that the base material on each side of a butt
splice will be from different mills, grades, or heats. When
this is the case, a separate transfer correction is required on
each plate at the weld to correct for differences in acoustic
properties. It is possible that the equipment settings may
differ for inspection of the same weld depending on which
side of the weld is being inspected.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
39 The research team recommends that additional requirements
be provided to account for the significant loss of amplitude
found when sound propagates through coarse grained NGIESW welds. These requirements include verifying the ampli­
tude and location of a 1.5 mm (0.06″) diameter SDH in a fullscale mockup of the weld. (Note, full-scale means simply full
thickness. The specimen need only be a few inches wide.) The
reflector would be required to be placed in a location that will
maximize the sound path traveling through the weld metal. This
recommendation will allow for an amplitude correction to be
determined on a case-by-case basis in order to account for the
high attenuation found when the sound beam passes through
the weld metal of some NGI-ESW welds. Recommendations
are also provided that NGI-ESW welds shall be inspected from
the outside of each fusion face. This recommendation, provided in the section that covers scanning coverage of welds,
will ensure that detection of fusion flaws will not depend on
sound passing through the entire weld volume.
3.5 Shear Wave Velocity
3.5.1 Experimental Testing
When inspecting a component using shear waves, sound
enters the material at some defined angle, usually ranging
between 45° and 70°. Conventional UT utilizes a single angle—
e.g., 70°—whereas PAUT utilizes a range of angles, typically
from 45° through 70°, or any specific group thereof (e.g., 55°
through 65°). An accurate estimate of the velocity through the
thickness is required since it is needed (using Snell’s Law) to
predict the refraction angle of the sound entering the steel. For
example, in conventional UT, when one uses a 70° wedge, the
geometry of the wedge is based on the assumed velocity of
0.127 in/µs (3,230 m/s) for shear waves to ensure the sound is
traveling through the steel at 70°. Unfortunately, even a very
small difference has a large impact on this calculation, especially for higher beam angles. According to research performed
in Japan and the CIVA simulation performed by the RT, differences as little as 2% or even less need to be accounted for,
as the sound will enter at a different angle than assumed. The
result is twofold. First, a given flaw will be mislocated since
the assumed geometry is not correct (e.g., the beam is not at
70°, but rather, something greater). Second, since the actual
sound path that is needed to maximize the reflection is different (and longer) and much of the sound energy is not actually hitting the reflector due to refraction into surface waves,
the resulting returning sound associated with a given reflector will have lower amplitude. From the perspective of the
technician, a given reflector will appear to be at the incorrect
location and to reflect less sound.
The most pronounced effects are observed at steeper incidence or sound beam angles, such as 60° and greater, where
a significant portion of the sound beam begins to reach the
critical angle at which the shear wave converts to a surface
wave (to be discussed in detail later). Note, when the term
“incidence angle” or “sound beam” is used, it is simply referring to primary angle—for example, 70°—though some of
the beam naturally spreads above and below, similar to how
a flashlight beam spreads out. The result is that some—or
in some cases, most—of the sound never actually enters the
plate as a shear wave but rather mode coverts and travels
along the surface of the plate at the steeper incidence angles.
This has been confirmed by both experimental testing and
CIVA simulations performed by the RT. Again, this is most
pronounced at beam angles greater than 60°. If a constant
index is used, there may be very little sound coverage in the
region of interest due to the beam entering at a different angle
than assumed, and if there is some sound reflected from a
flaw, the amplitude is reduced, sometimes significantly.
To investigate this further, the acoustic shear wave velocity
in a number of TMCP HPS 70W plates was measured using
an Electro Magnetic Acoustic Transducer (EMAT). Using a
shear wave, acoustic velocity was measured in the rolled and
transverse to the rolled direction. The acoustic velocities and
anisotropic ratios are listed in Table 31. The anisotropic ratio
is the measured difference between the acoustic velocities in
the rolled direction (RD) and transverse to rolled (⊥RD) direction. The amount of anisotropy is defined by the anisotropic
ratio, which is the ratio of the velocity in the rolled direction
to the transverse to rolled direction, as shown below.
 Velocity in RD 
Anisotropic Ratio (%) = 
− 1 × 100
 Velocity in ⊥ RD 
Table 31. Shear wave acoustic velocities of steel specimens.
ID
70 (QT)
101 (QT)
TMCP 1
TMCP 2
TMCP 3
Acoustic Velocity (in/µsec)
Transverse to Rolled
Rolled Direction
Direction
0.1271
0.1272
0.1272
0.1274
0.1328
0.1266
0.1304
0.1241
0.1293
0.1255
Anisotropic
Ratio (%)
0.05
0.15
4.95
5.07
3.09
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
40
The QT specimens—specimens 70 and 101—have a very
low anisotropic ratio. In comparison, it is clear that all of the
TMCP specimens demonstrate high anisotropic ratio. Plates
with a high anisotropic ratio(1–2% or greater) are considered
an acoustically anisotropic material [45]. All TMCP plates
tested demonstrated significant acoustic anisotropy. In other
words, there is directional variation in velocity in the as-rolled
and cross-rolled directions. The average velocity through the
thickness differed by as much as 5%, which, as noted, at this
level has a significant effect on the refraction angle.
After the observation, an additional literature review was
performed to establish if other researchers had observed this
same behavior in TMCP plates. Previous research in Japan
has documented this effect for TMCP plates [45, 46]. The
Japanese JIS Z 3060 UT code [13] specifies that the shear
wave velocity be measured in the direction in which the
inspection will occur in the test object and compared to the
calibration block. Depending on the ratio of the shear wave
velocity in the calibration block and test specimen, either a
new calibration block is required with a velocity that matches
the test object more closely or restrictions are placed on the
incidence angle that may be used in the inspection. All three
TMCP specimens used during this study would have exceeded
the limits that would have corrective action as acoustically
anisotropic using the Japanese criteria.
It is noteworthy that this is also an issue for conventional
UT, as the 70° probe is the prescribed transducer for most
applications. As will be shown for some TMCP plates, due
to the refraction issues, the 70° probe can be quite ineffective in detecting flaws. This is concerning and, though out of
the scope of this project, this will also need to be addressed
in the current AWS D1.5. Other specifications require the
acoustic velocity of the plate be measured and used to determine the true angle of refraction with respect to the assumed
scanning angle. When a significant difference between measured and assumed velocity is observed (1% or greater is
considered significant), the scanning angle is often limited
to no greater than 60°. For brevity, the specific details of the
experimental testing program which measured the shear
wave velocity of base metal are given in Appendix F.
Additional velocity measurements were made on other
grades of base metals commonly used in bridges using
a normal incidence shear wave probe, shown in Table 32.
Machined specimens were not fabricated from these plates,
so a correlation of the shear wave velocity to the amplitude
response of a SDH is not provided in these cases. These data
are simply provided to aid in determining whether other
grades are susceptible to velocity variations. The A709-HPS
50W and A709-HPS 70W specimens both had a shear wave
velocity in the rolled direction that would differ from a standard velocity of 0.127 in/µs by more than 1%, along with anisotropic ratios greater than 1%. Therefore, both of these steels
would be susceptible to calibration issues due to mismatch of
the calibration block and test object shear wave velocity as well
as beam splitting when testing along an oblique orientation to
the rolled direction due to acoustic anisotropy.
3.5.2 AWS Recommendation
for Shear Wave Velocity
3.5.2.1 Rolled Direction or Transverse
to Rolled Direction
Differences in measured shear wave velocity for common
grades of base metals were noted during the experimental testing. For TMCP processed steels tested, the velocity
increased in the rolled direction and decreased in the transverse to rolled direction. These materials are considered
acoustically anisotropic. Since the incidence angle is related
to the ratio of the wedge velocity to the steel velocity by
Snell’s Law, differences in the incidence angle result in errors
in properly locating flaws as well as affecting the amplitude
reflected from the indications (i.e., a reduction in amplitude).
Experimental testing found that the amplitude was significantly decreased at high incidence angles (60° or greater)
when scanning was performed in a direction of increased
velocity. CIVA models were developed that represented one of
the TMCP processed steel samples from the experimental testing with a shear wave velocity of 0.133 in/µs (4.5% increase
compared to the calibration block). The CIVA models showed
that the difference in amplitude of the TMCP steel to the 1018
calibration block would be ∼2 dB at a 7″ depth with 45° incidence angle, would be ∼2 dB at a 1″ depth and ∼4 dB at a
7″ depth with 60° incidence angle, and would exceeded 2 dB
Table 32. Shear wave acoustic velocities of other bridge base materials.
Grade
A709-50
A709-50W
A709-HPS 50W
A709-HPS 70W
(TMCP)
Acoustic Velocity (in/µsec)
Transverse to Rolled
Rolled Direction
Direction
0.1281
0.1280
0.1279
0.1275
0.1287
0.1270
0.1294
0.1262
Anisotropic Ratio (%)
0. 05
0. 36
1.29
2.54
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
41 for all sound paths with the 70° incidence angle. Therefore,
while limiting the incidence angle range to 40°–60° will lessen
the impact from the changes in the shear wave velocity, it will
not limit the amplitude differences to 2 dB or less over all
possible sound paths.
Additional CIVA analysis was performed in order to identify
the limiting velocity difference compared to the standard steel
velocity of 0.127 in/µs (3,230 m/s) which would result in an
amplitude difference of 2 dB or less over a sound path cover­
ing 7″ depth for the 60° and 70° incidence angles. A velocity
increase of 2.5% compared to the standard velocity resulted
in an amplitude difference of 2 dB or less across the 40°–60°
incidence angles. Therefore, assuming a shear wave velocity
of 0.127–0.128 in/µs for the calibration block, which is typical for most steels, the 40°–60° incidence angles should be
appropriate for plates with velocity of 0.130–0.131 in/µs.
Two of the three TMCP samples from the experimental testing had a shear wave velocity of 0.130 in/µs or less, and the
amplitude difference for these two samples compared to the
Grade 50 sample were 2 dB or less at the 60° incidence angle.
This correlates well with the CIVA results regarding a recommended limit of 2.5% velocity difference for inspection up to
a 60° incidence angle. Limiting the incidence angle range to
40°–60° for inspection of TMCP processed steels will result
in much less amplitude error than using 40°–70° and may
limit the amplitude error to 2 dB or less for many TMCP processed steels, however there may be certain heats of TMCP
processed steels where the amplitude loss may exceed 2 dB at
the 60° incidence angle.
A velocity increase of 1.0% compared to the calibration
block resulted in an amplitude difference of 2 dB or less across
the 40°–70° incidence angles. Assuming a shear wave velocity
of 0.127–0.128 in/µs, the 40°–70° incidence angles should be
appropriate for plates with velocity of 0.128–0.129 in/µs. One
of the three TMCP samples from the experimental testing
had a shear wave velocity of 0.1293 in/µs, and the amplitude
difference for this sample compared to the Grade 50 sample
(measured shear wave velocity of 0.1274 in/µs) was 2–4 dB at
the 70° incidence angle. The velocity of this TMCP sample
was 1.5% greater than the Grade 50 sample, so a slight increase
in sound loss exceeding 2 dB is expected at the 70° incidence
angle. While a 1.0% limit on the velocity difference between
the calibration block and the test object seems reasonable to
limit the amplitude difference to 2 dB or less over the 40°–70°
incidence angle range, the actual amplitude difference may
exceed 2 dB in some cases, especially when also including any
differences in material attenuation, as discussed previously in
Section 3.4.
A marked-up version of Annex K and associated commentary with the recommended changes is given in Appendix G.
The research team is recommending that the acoustic
properties of the test object and calibration block be verified to
be within certain tolerances. It is recommended that the shear
wave velocity in the direction of sound propagation in the
test object and the calibration block be required to be within
±2.5% of each other. When the difference exceeds this amount,
it is recommended that a new calibration block with a velocity within 2.5% of the test object be fabricated or otherwise
acquired. When the difference in shear wave velocity exceeds
±1%, the research team recommends that the incidence angle
be limited from 40°–60°. If the velocity of the test object and
calibration block is measured and found to be within ±1% of
each other, a 40°–70° incidence angle range may be used. Since
most steels have a velocity of 0.127–0.128 in/µs, these steels
would all be able to use the same calibration block, assuming
that attenuation is properly accounted for.
JIS Z 3060:2015 [13] includes three different methods to
directly or indirectly measure the shear wave velocity of the
test object and calibration block:
1. Fabricate a block with a reference reflector from the test
object material. Calculate the incidence angle of the test
object and the calibration block by using measurements of
the physical distance of the reflector from the index point
and the depth of the reflector.
2. Perform pitch–catch measurement on the test object and
calibration block in a single-V path, and calculate the incidence angle by using measurements of the physical distance
between the transducer index points and the thickness of
the plate.
3. Directly measure the shear wave velocity of the plate by
using a normal incidence shear wave probe to measure
backwall signals. The shear wave velocity of the calibration block and test object can either be directly computed
using a calibration feature with successive shear wave
backwall skip signals (similar to velocity measurements
using a normal incidence longitudinal wave probe), or
the velocity ratio between the calibration block and test
object can be computed using the known thickness and
measured shear wave backwall signals of each plate.
If the incidence angle is measured, the velocity ratio can be
calculated using Snell’s Law as follows:
sin θtest object Vtest object
=
sin θcal block
Vcal block
Assuming that the calibration block incidence angles are
truly 60° and 70° (i.e., velocity of calibration block matches
the assumed velocity), the incidence angles measured in the
test object are 62.6° and 74.4° for a 2.5% increase in velocity
and 61° and 71.6° for a 1.0% increase in velocity.
All of these methods rely on the orientation of the probe
used to measure the velocity or incidence angle matching the
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
42
calibration blocks for TMCP steels must have the correct
scanning surface and scanning direction (e.g., throughthickness and rolled direction) compared with the test object,
as shown on the right side of Figure 11.
3.5.2.2 Oblique to Rolled Direction
Scanning
Surface Rotated
Figure 11. Calibration block rotation.
orientation of the probe that will be used during scanning. For
instance, if the PAUT line scanning will be performed with the
probe oriented along the rolled direction, then the measurement of the shear wave velocity or incidence angle shall be
performed in the same direction. As an example, checking the
depth and index distance from the corner trap signal, which
is produced from scanning the edge of the plate transverse
to the rolled direction, cannot be used to verify the velocity
along the rolled direction. Fabrication of a calibration block is
also sensitive to the orientation of the plate in relation to the
orientation of the scanning direction.
In fact, the direction of sound propagation in regards to
the through-thickness direction is also critical. For instance,
typically calibration blocks are cut from material and flipped
so that the through-thickness direction forms the width of
the calibration block, as shown for the block on the left side
of Figure 11. This is performed in order to provide for greater
depths of the SDH reflectors and avoid the need for very thick
plates. Due to the complex and layered grain structure found
in TMCP steels, the velocity at the surface of the plate will
differ from the velocity in the middle of the plate. Therefore,
1st skip
Since the TMCP steels have very different velocity in the
rolled direction and the transverse to the rolled direction,
they are considered acoustic anisotropic materials. When
scanning is performed oblique to the rolled direction (i.e.,
not parallel or perpendicular), the beam splits into two waves
traveling at different velocities, as shown in Figure 12 for the
normal incidence shear wave probe. Since the beam is split,
the amplitude is effectively halved, and it is possible for two
indications to appear on the screen, as shown in Figure 13.
Experimental testing was performed on Specimen TMCP
2 plate, cut at a 45° orientation to the rolled direction with
1⁄16″ dia. SDHs at 0.6″ and 1.0″ depth. The machined specimen dimensions are similar to the other TMCP specimens.
The amplitude difference for the 45°, 60°, and 70° incidence
angles on the 45° orientation to rolled direction TMCP
block compared to the Grade 50 block is shown in Figure 14
for the 2.25 MHz Zetec PAUT probe (AXL-2.25 MHz with
AXL-55SW wedge with 16 active elements) and an active
aperture of 16 mm × 20 mm. The Grade 50 block traditionally served as reference for the other TMCP testing.
The results from testing using the same heat of steel but
in the rolled and transverse to rolled orientation are shown
in Figure 15. This testing was performed using a 2.25 MHz
conventional UT probe, which had a similar aperture to the
PAUT probe. Comparing these figures, the 45° orientation to
the rolled direction block had more loss of amplitude at the
45° and 60° incidence angles than the rolled and transverse
to rolled blocks. This is likely due to the beam-splitting effect
for the 45° orientation block. The loss of amplitude in the 45°
orientation block at the 70° incidence angle was comparable
2nd skip
3rd skip
Figure 12. Normal incidence shear wave probe at 45ç angle.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
43 Figure 13. 0.6” deep SDH in 45ç TMCP plate.
16.0
14.0
12.0
10.0
dB
8.0
0.6"
6.0
4.0
2.0
0.0
1.0"
0.6" 1.0"
0.6" 1.0"
Reference
-2.0
-4.0
45ᵒ
60ᵒ
70ᵒ
Search Angle
Figure 14. 45ç orientation TMCP plate amplitude difference
for 2.25 MHz PAUT probe.
45° 60° 70° TMCP 2
Roll
x x + -roll
16.0
14.0
12.0
0.6"
10.0
dB
8.0
1.0"
6.0
4.0
2.0
0.0
0.6" 1.0"
0.6"
1.0"
-2.0
-4.0
45ᵒ
60ᵒ
70ᵒ
Search Angle
Figure 15. Rolled and transverse to rolled orientation TMCP plate
amplitude difference for 2.25 MHz conventional UT.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
44
to the results of the rolled direction for the 1″ deep hole but
was lower than the rolled direction for the 0.6″ hole.
Based on these results, the research team recommends additional requirements for acoustically anisotropic materials. It
recommends that these requirements be placed on materials
that have an anisotropic ratio (ratio of velocity in the rolled
and transverse to rolled direction) of 1% or greater. These recommendations include a requirement that the incidence angle
range be limited to 40°–60° for scanning of acoustic anisotropic plates (e.g., TMCP) at an oblique orientation to the rolled
direction. It is also recommended that 4 dB required to be
added to the reference sensitivity to account for the sound loss
due to the beam splitting.
While normal incidence shear wave probes can be used
to quickly ascertain whether a plate has acoustic anisotropic
behavior (i.e., the velocity is different in the rolled and transverse to rolled directions) by rotating the probe from polarization in the rolled direction to the transverse to rolled direction,
caution may be necessary to use it for measuring the velocity
in the oblique orientation. For scanning along oblique orientations in acoustic anisotropic materials, JIS Z 3060 requires
that the velocity be accounted for by measuring the incidence
angle using an angle beam probe(s) rather than directly measuring the velocity with the normal incidence shear wave
probe. While the technical reason for this requirement is not
explained in JIS Z 3060, it is likely due to the fact that two different measurements are possible with the normal incidence
shear wave probe, which could result in miscalculations. Measurement of the incidence angle through either fabrication of
a reference standard or using the pitch–catch method results
in measurement of the true refraction angle at the maximum
amplitude indication. There may still be two indications
apparent on the screen, but the maximum amplitude indication would be used for measurement of the incidence angle.
The JIS code does not require additional amplitude be added
for scanning in the oblique orientation, but it does require
that the calibration standard be acoustically equivalent to the
test object with velocity within ±2% and sensitivity correction
within ±2 dB compared to the test block along with additional
limits on the incidence angle depending on the ratio of the
velocity of the test object to the calibration block in the orientation of the scanning direction.
Based on the experimental results, additional requirements
are recommended for acoustically anisotropic materials. It is
recommended that all materials with an anisotropic ratio
(a ratio of velocity in the rolled and transverse to rolled direction) of 1% or greater be defined as acoustically anisotropic
in Annex K. Based on the experimental test results, additional
requirements on the incidence angle and addition of amplitude are recommended for scanning of acoustically anisotropic materials in the oblique orientation with respect to
the rolled direction. It is recommended to limit the incidence
angle to 40°–60° for scanning of acoustic anisotropic plates
at an oblique orientation to the rolled direction. It is also recommended that 4 dB be added to the reference sensitivity to
account for the sound loss due to the beam splitting.
These requirements do not require the velocity or beam
incidence angle be measured in the oblique orientation. Therefore, while guidance is provided in the commentary that caution must be taken when measuring the velocity of the plate
in the oblique orientation using a normal incidence shear
wave probe, this measurement is not necessary. Rather, acoustic anisotropic materials are identified using the shear wave
velocity measurements in the rolled and transverse to rolled
directions and checking against the 1% limit on difference.
3.6 Calibration Blocks
While correction for differences in attenuation and velocity
between the calibration block and the test object may be provided through the use of certain probes and incidence angles or
through the use of a transfer correction, it can also be provided
by using a calibration block that is acoustically equivalent to
the test object. Thus, fabrication of additional calibration
blocks that are acoustically equivalent to the various steels
commonly used in a shop may be prudent. As discussed previously, the orientation and scanning face of the calibration
block is also critical for acoustic anisotropic material such as
TMCP processed steel. For this reason, guidance is provided
on recommendations for proper calibration block design. Due
to the necessary sensitivity for flaw detection and rejection discussed in the following sections, it is recommended that the
1.5 mm (0.06″) diameter SDH reference standard reflector be
used for calibration and setting reference amplitude. This is
the same reference standard reflector that is currently being
used in AWS Annex K; no changes to the code in regards to the
reference standard reflector are necessary.
Calibration blocks can be machined from a strip of steel
removed from a plate with 0.06″ (1.5 mm) diameter SDHs
drilled through the width. In order to provide enough sound
paths for TCG calibration, it is recommended that one SDH
be placed near the surface of the plate and the other placed
either in the center or third-point, depending on the plate
thickness. This provides for many possible TCG points, as
shown in Figure 16. It is recommended that the current minimum of 3-point TCG be carried forward in the new version
of Annex K, but nine or more TCG points are possible from
this block. The hole near the surface should be placed at least
0.2″ away from the surface in order to distinguish the first leg
indication from the second leg indication. The hole 0.2″ away
(minimum) from the surface should not be scanned with a
skip off of the near surface since the small ligament between
the edge of the hole and surface of the plate can result in
increased amplitude. This is similar to a corner trap from a
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
45 Figure 16. Possible TCG scanning positions.
surface breaking flaw. CIVA analysis was performed to determine the minimum ligament for the SDH from the surface
of the plate in order to use it for skipping off of the backwall.
A 0.5″ minimum depth from the surface of the plate should
be provided in order to skip from the near surface. It is
recommended that this limit be included in the calibration
block geometry requirements in Annex K.
The width of the block should be large enough to
accommodate the beam spread without funneling the sound
beam along the calibration block at long sound paths. The
JIS Z 3060 code includes the following equation for estimating a suggested minimum width of the calibration block:
W >2×l×S D
Where:
W: width of the calibration block
l: wavelength
S: maximum sound path to be used
D: width of the transducer
Using typical values in this equation—for example, a
11.7″ sound path, which would represent a 4″ depth at 70°
incidence angle—the minimum width of the block would
be 2.1″ for a 2.25 MHz probe with a 16 mm width and 1.7″
for either a 2.25 MHz probe with a 20 mm width or a 5 MHz
probe with a 10 mm width. Therefore, a 2″ width will likely
be adequate for most blocks, although a narrower block
may be appropriate if it is used over shorter sound paths.
It is recommended that Annex K include the requirement
that the calibration block be of sufficient width to allow for
adequate beam spread at the longest sound path used for
calibration. A marked-up version of Annex K and associated commentary with the recommended changes is given
in Appendix G. The JIS code equation is also provided in
the commentary as a recommended minimum width of the
calibration block.
Finally, the length of the block should be adequate to
accommodate multiple skips for TCG calibration and pitch–
catch comparison to the test object. It is recommended that
the spacing of the two holes be wide enough to provide for
the double-V-path in the pitch–catch setup for the 70° inci-
dence angle. This will also provide adequate clearance for
multiple skips for the TCG calibration.
The holes should be spaced far enough from the end of the
plate to limit the corner trap signal from affecting the TCG
calibration. For instance, as a PAUT probe is swept over the
SDH for TCG calibration, the high incidence angle can hit
the corner trap before the low incidence angle has been calibrated. If the corner trap signal has greater amplitude than
the SDH, the corner trap amplitude will incorrectly be used
for the TCG calibration rather than the SDH amplitude.
There are two ways to avoid this issue: (1) space the SDH
far enough from the end of the plate so that the corner trap
is not reached at the high incidence angles, or (2) separate
the incidence angle range into multiple groups and perform
TCG calibration on each group separately. For instance, the
40°–70° incidence angle range could be split into a 40°–60°
range and a 60°–70° range and each range swept over the SDH
independently. The TCG for each angle range is then combined within the instrument software to provide for a single
TCG covering the entire angle range. This is particularly an
issue for long sound paths since the coverage between the 40°
and the 70° beams is quite large.
Based on preliminary CIVA analysis, it is recommended
that the SDHs should be spaced at least 5″ from the end of
the plate with a 1″ plate thickness, although the TCG may still
require splitting into two groups for long sound paths. The
recommended dimensions for a 1″ thick calibration block are
shown in Figure 17, but the SDH spacing and plate width may
be modified based on the specific probe used in the inspection. The dimensions for the plate length, width, SDH depth,
spacing, and placement may all be different for different plate
thicknesses. For instance, the recommended dimensions for
a 2″ thick calibration block are shown in Figure 18. This plate
would also require the TCG be split into two groups for long
sound paths, such as the third leg 4.5″ depth TCG point. The
2.5″ width was determined by using the 70° incidence angle
for a 2.25 MHz 16 mm wide probe (equivalent to 5 MHz
8 mm wide probe) for the third leg 4.5″ depth TCG point
since this point would cover first and second leg scanning
for 2″ thick material. It should be noted that machining a
0.06″ dia. SDH through 2.5″ thick material may be difficult
due to the short length of available drill bits. Recommended
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
46
Figure 17. Recommended calibration block for 1” thick plate.
placement of ultrasonic transducers for transfer correction
and TCG calibration are shown in Figure 9 and Figure 16,
respectively.
Finally, the research team recommends that Annex K require
the calibration block to be similar in temperature to the test
object when calibration is performed. During the attenuation
experimental testing, differences in attenuation measurements
were noted when measurements were taken at different
temperatures. It is recommended that the temperature limits
from the AWS D1.1 PAUT proposal of ±25°F be included in
AWS D1.5 Annex K.
3.7 Amplitude Limit for Rejection
of Flaws
3.7.1 CIVA Modeling
Modifications to the AWS D1.5 Annex K acceptance criteria
were developed in order to improve detection and rejection
of critical weld flaws. The approach to the inspection procedure is to use line scanning for detection of weld flaws and
follow-up raster scanning to evaluate rejection of weld flaws.
The amplitude to be compared to the acceptance criteria
limits will be the maximum amplitude measured during raster scanning. Raster scanning will involve scanning the indication over the entire incidence angle range from both sides
of the weld while also rotating the probe. Compared with line
scanning, raster scanning will maximize the amplitude as it
will involve movement of the probe to account for unknown
parameters such as flaw tilt, flaw skew, and flaw location.
In order to develop the amplitude limit for flaw rejection,
CIVA analysis was performed using a 2.25 MHz PAUT probe
with an active aperture of 16 mm by 16 mm and an angular range of 45°–70°. Using CIVA, the maximum amplitude
that would be found during typical raster scanning with TCG
was determined. The CIVA analysis was performed on flaws
similar in size to the critical planar and volumetric flaws. The
amplitude from the indications was referenced to the amplitude of a 1.5 mm (0.06″) diameter SDH at a similar sound
path, with positive amplitude having a higher response than
the SDH and negative amplitude having a smaller response
than the SDH. Therefore, positive results are more severe than
negative results. Flaw tilt of planar flaws were evaluated at 0°,
±5°, ±30°, and ±45°. The least positive maximum amplitude
measured over all of the evaluated flaw tilts is the control used
to reject critical flaws regardless of their tilt. In other words,
the smallest possible amplitude that would result from raster
scanning, regardless of the actual flaw tilt, forms the basis of
the rejection limit.
Since the critical flaw size depends on the stress range, stress
concentration from thickness transitions, and the throughthickness location of the flaw, different critical flaw sizes are
possible depending upon these inputs. The critical flaw sizes
for surface and embedded flaws in equal thickness and thickness transition welds were evaluated for 4 ksi and 8 ksi stress
ranges. Plate thicknesses of 0.5″ and 2″ were used to account
for variations due to the probe near field and natural beam
shape. The results of the parametric CIVA modeling for the
controlling maximum amplitude compared to the 1.5 mm
(0.06″) diameter SDH are shown in Table 33 and Table 34 for
critical planar and volumetric flaws, respectively.
The results of this analysis are further summarized in
Table 35 by combining the results of embedded and surface
flaws for the same stress range and weld type (i.e., equal thick-
Figure 18. Recommended calibration block for 2” thick plate.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
47 Table 33. CIVA results of controlling maximum amplitude for critical
planar flaws.
Critical Planar
Flaw
Stress
Range
Flaw
Location
Embedded
4 ksi
Surface
Embedded
8 ksi
Surface
Weld
Type
Comparable CIVA Analysis
Height Length Height
(in)
(in)
(in)
Equal
Transition
Equal
Transition
Equal
Transition
Equal
Transition
0.20
0.17
0.10
0.06
0.14
0.08
0.06
0.02
0.20
0.17
0.13
0.08
0.14
0.08
0.08
0.03
0.20
0.15
0.10
0.05
0.15
0.10
0.05
0.025
Length
(in)
Plate
Thickness
(in)
0.20
0.15
0.15
0.10
0.15
0.10
0.10
0.025
0.5/2
0.5/2
0.5/2
0.5/2
0.5/2
0.5/2
0.5/2
0.5
Controlling
Maximum
Amplitude
(dB)
-6
-8
-2
-10
-8
-10
-10
-38
Table 34. CIVA results of controlling maximum amplitude for critical
volumetric flaws.
Stress
Range
Flaw
Location
Weld
Type
Equal
Transition
Equal
Surface
Transition
Equal
Embedded
Transition
Equal
Surface
Transition
Embedded
4 ksi
8 ksi
Critical
Volumetric
Flaw
Comparable CIVA Analysis
Height Length Diameter
(in)
(in)
(in)
0.22
0.24
0.23
0.08
0.14
0.08
0.06
0.02
0.41
0.24
0.31
0.11
0.14
0.08
0.08
0.03
0.25
0.25
0.25
0.08
0.125
0.08
0.08
0.03
Plate
Thickness
(in)
0.5/2
0.5/2
0.5/2
0.5/2
0.5/2
0.5/2
0.5/2
0.5
Controlling
Maximum
Amplitude
(dB)
-7
-7
-7
-13
-13
-18
-13
-18*
*Result may not be valid since standard analytical CIVA model is not valid due to small flaw size relative
to the wavelength
ness or thickness transition). This forms the basis of possible
acceptance criteria amplitude limits to be applied to the raster
scanning results. The acceptance criteria amplitude limit
varies from –6 dB (indication amplitude 6 dB below reference amplitude) for critical planar flaws in equal thickness
welds under 4 ksi stress range to –18 dB for critical embedded
volumetric flaws in thickness transition welds under 8 ksi stress
range. It should be noted that critical surface planar and voluTable 35. Raster scanning acceptance
criteria limits from CIVA.
Stress
Range
4 ksi
4 ksi
8 ksi
8 ksi
Weld Type
Equal Thickness
Thickness Transition
Equal Thickness
Thickness Transition
(Embedded Flaws Only)
Planar
Flaws
-6 dB
-10 dB
-10 dB
Volumetric
Flaws
-7 dB
-13 dB
-13 dB
-10 dB
-18 dB
metric flaws in thickness transition welds under 8 ksi stress
range are not included in this table since the critical flaw size
according to fracture mechanics was so small that the standard analytical CIVA models are not valid (approximately
half the wavelength of the 2.25 MHz probe).
3.7.2 Comparison to Current Amplitudebased Acceptance Criteria
Figure 19 compares the acceptance criteria amplitude
limits from the CIVA analysis in Table 35 to the current acceptance criteria in AWS D1.5 Annex K for flaws in the middle
half of tension welds. In this figure, the maximum amplitude
from the indication compared to the 1.5 mm (0.06″) diameter SDH is on the vertical axis and the indication length is
along the horizontal axis. Combinations of amplitude and
length that fall below the line would be accepted while those
above the line are rejected. For instance, AWS Annex K would
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
Maximum Amplitude (0.06 inch dia. SDH for Reference)
(dB)
48
10
Class A Limit
Rejectable
5
Class B Limit
0
CIVA -6 dB
Class C Limit
CIVA -7 dB
-5
CIVA-10 dB
-10
CIVA-13 dB
More
Conservative
-15
-20
Acceptable
0
1
2
3
Indication Length (in)
4
5
Figure 19. Comparison of CIVA results to AWS D1.5 Annex K acceptance
criteria.
accept an indication that was up to 5 dB above reference [i.e.,
5 dB greater amplitude than the 1.5 mm (0.06″) diameter
SDH] as long as the indication length is 0.5″ or less. Also,
AWS Annex K would accept any indication which is more
than 6 dB below reference [i.e., amplitude of –6 dB or less
compared with the 1.5 mm (0.06″) diameter SDH] regardless
of the indication length.
It is apparent from Figure 19 that the CIVA amplitude limit
results are more conservative than the AWS Annex K acceptance criteria limits since the AWS Annex K line is above the
CIVA results. For instance, Annex K acceptance criteria has
higher amplitude limits for Class A and B flaws than any of
the CIVA results. Therefore, there is more area for indications
to fall under the Annex K line in the acceptable zone rather
than the CIVA results. Once again, since Annex K also allows
for evaluation of amplitude based on the initial line scan, it is
unlikely that the amplitude used to compare to the acceptance
criteria will be the maximum flaw amplitude due to possible
flaw tilt, skew, and location in relation to the probe. The lack
of maximizing the flaw amplitude, which is compared to the
acceptance criteria for Annex K, would compound the differences between an evaluation using Annex K and using raster
scanning with the CIVA-determined limits.
In order to compare the proposed amplitude limits from
CIVA to the conventional UT amplitude limits, the conventional UT amplitude limits must be inverted from positive
to negative integers. This is due to the fact that the conventional UT and PAUT have different approaches to obtain
the indication amplitude. Conventional UT, according to
AWS D1.5, requires modifying the instrument gain in order
to force the indication amplitude to match the reference
amplitude of the 1.5 mm (0.06″) diameter SDH. Indications
with high amplitude result in a low (or negative) indication
rating since instrument gain is removed in order to bring
the indication amplitude down to reference amplitude. For
instance, in Figure 20 (left), the indication exceeds the reference amplitude at reference gain; this indication has greater
amplitude than the reference standard. In Figure 20 (right),
4 dB gain was removed from the instrument to bring the
indication amplitude to the level of the reference amplitude. Assuming that the attenuation factor would be zero
in this case, this indication would have an indication
rating of –4 dB. This is the opposite of PAUT testing where
the indications with high amplitude result in a more positive
reported amplitude compared to the reference amplitude. For
instance, PAUT inspection of the indication in Figure 20
would have a reported amplitude of +4 dB compared to the
reference amplitude.
Figure 21 compares the acceptance criteria amplitude limits
from the CIVA analysis in Table 35 to the current acceptance
criteria in AWS D1.5 Clause 6 conventional UT for flaws in the
middle half of the tension welds after inverting the positive to
negative values for the conventional UT tables. Conventional
UT in Clause 6 includes separate criteria depending on the
plate thickness and incidence angle used in the inspection.
Rather than having one line like Annex K, there are eight different lines for conventional UT acceptance limits. It is apparent that the conventional UT acceptance criteria are more
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
49 Figure 20. Conventional UT amplitude measurement.
on the historical performance of workmanship-based acceptance criteria but rather on FFS.
The 2018 edition of the CSA W59 code [11] includes an
alternative acceptance criteria based on TCG rather than the
fixed attenuation approach. This acceptance criteria was based
on the CSA fixed attenuation conventional UT tables, which
are similar to the AWS D1.5 Clause 6 conventional UT tables.
The CSA W59 alternative TCG acceptance criteria may be
applied to conventional UT or manual raster scanned PAUT.
Encoded PAUT is specifically noted as an alternative ultrasonic system which is subject to a written agreement between
the engineer and contractor along with development of an
appropriate scanning procedure and acceptance criteria. The
CSA TCG acceptance criteria were derived by combining the
Maximum Amplitude (0.06 inch dia. SDH for Reference) (dB)
conservative than AWS D1.5 Annex K since the CIVA results
from –6 dB to –13 dB correlate to the conventional UT limits
while they are in the acceptable range for Annex K.
Although it may seem that the CIVA results are overly
conservative when compared to Annex K, they correlate well
to the conventional UT acceptance criteria, which appears
to have provided good historical performance when used
for UT inspection. It is also important to remember that the
amplitude limits from CIVA were derived completely independent of the Clause 6 conventional UT acceptance criteria
by modeling critical flaws derived from fracture mechanics.
While the similarities of the CIVA results to traditional
acceptance criteria help to bolster confidence in the newly
derived amplitude limits, these limits are not based solely
10
Rejectable
5
CIVA -6 dB
CIVA -7 dB
Class A Limit
CIVA-10 dB
CIVA-13 dB
Class B Limit
0
CIVA -18 dB
Class C Limit
Conventional UT 5/16"-3/4" PL 70 Deg
Conventional UT >3/4"-1.5" PL 70 Deg
-5
Conventional UT >1.5"-2.5" PL 70 Deg
Conventional UT >1.5"-2.5" PL 60 Deg
Conventional UT >1.5"-2.5" PL 45 Deg
-10
Conventional UT >2.5"-4" PL 70 Deg
Conventional UT >2.5"-4" PL 60 Deg
Conventional UT >2.5"-4" PL 45 Deg
-15
-20
More
Conservative
Acceptable
0
1
2
3
Indication Length (in)
4
5
Figure 21. Comparison of CIVA results to AWS D1.5 conventional UT
acceptance criteria.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
50
amplitude limits for the various incidence angles in the fixed
attenuation tables along with correction for the difference
in true attenuation compared to the fixed attenuation model
[14–16]. By taking this approach, the CSA code attempts to
recreate the same level of quality using the TCG acceptance
criteria as the traditional fixed attenuation acceptance criteria.
Figure 22 compares the TCG acceptance criteria in CSA
W59-18 to the CIVA results for flaws in the middle half of the
tension welds. Since the CSA TCG acceptance criteria depends
on the plate thickness, there are three different criteria plotted
in this figure. It can be seen that the CIVA results of –7 dB
and –10 dB cross the CSA acceptance criteria and, therefore,
are quite similar. The CIVA result of –6 dB would have been
rejectable regardless of the indication length, according to the
CSA TCG acceptance criteria for all of the plate thicknesses;
therefore, the –6 dB limit would be slightly less conservative
to the CSA TCG limits. On the other hand, the CIVA results
of –13 dB and –18 dB would be acceptable regardless of the
indication length and therefore would be more conservative.
3.7.3 Verification Testing of Weld
Flaw Samples
Initial verification of the CIVA analysis was performed by
rescanning the round robin specimens with two different
PAUT probes and using a standard AISI 1018 steel calibration block with 1.5 mm (0.06″) diameter SDHs for TCG.
Testing was performed using a Zetec Topaz 16 with a Zetec
AXL-2.25 MHz PAUT probe with 16 active elements (active
aperture of 16 mm by 20 mm) and with a Zetec AM-5 MHz
PAUT probe with 16 active elements (active aperture of
9.6 mm by 10 mm). In general, it was found that the amplitude from the flaws was greater using the 5 MHz probe as
compared with the 2.25 MHz probe. This is not surprising
since the wavelength is smaller for the 5 MHz probe and it is
more sensitive to small flaws.
This testing found that the lowest maximum amplitude
from a planar flaw that would be rejectable according to
fracture mechanics (excluding the extremely small critical
size calculated for surface flaws in thickness transition welds
under 8 ksi) was +3.6 dB for a vertical crack which was
0.17″ high by 0.40″ long. Therefore, this flaw would have
been rejectable according to the amplitude limits found
during the CIVA analysis.
The round robin specimens also included some very small
planar flaws which would only be rejectable according to
fracture mechanics as a surface flaw in thickness transition
weld under 8 ksi. The lowest maximum amplitude from
these flaws was –5.3 dB for a surface breaking crack which
was 0.02″ high by 0.04″ long. This flaw would also have been
rejectable according to the amplitude limits found during the
CIVA analysis.
The lowest maximum amplitude from a volumetric flaw
was –13 dB for a near surface group of porosity, which was
0.09″ high by 3.31″ long with a maximum pore diameter
measured with RT of 0.05″. This pore diameter is similar to
Maximum Amplitude (0.06 inch dia. SDH for Reference) (dB)
10
Rejectable
5
0
CIVA -6 dB
Class A Limit
CIVA -7 dB
Class B Limit
-5
CIVA-10 dB
CIVA-13 dB
CIVA -18 dB
Class C Limit
CSA W59-18 5/16"1.5" PL
-10
CSA W59-18 >1.5"-2.5" PL
CSA W59-18 >2.5"-4" PL
-15
-20
More
Conservative
Acceptable
0
1
2
3
Indication length (in)
4
5
Figure 22. Comparison of CIVA results to CSA W59 TCG acceptance criteria.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
51 the critical volumetric flaw used in the CIVA analysis for the
4 ksi thickness transition and 8 ksi equal thickness cases. This
flaw would have been rejectable according to the amplitude
limits found during the CIVA analysis for these cases. Current
RT acceptance criteria would have also rejected this flaw.
Final verification of the CIVA analysis was performed by
an ASNT Level III UT/Level II PAUT technician. Round robin
specimens with low amplitude indications were utilized for
this testing along with additional flawed weld specimens that
were acquired specifically for this testing. These specimens
were acquired because they included known weld flaws that
were small in size relative to many of the flaws included in the
round robin specimens.
The final verification testing involved line scanning the
samples for flaw detection followed by raster scanning for
evaluation of acceptance. As the flaw detection amplitude
limit, which would require follow-up raster scanning, was
unknown at the time (explained further in the follow-
ing section), all indications with an amplitude greater than
–20 dB during the initial line scanning were further investigated with raster scanning to determine the maximum amplitude. Testing was performed with an Olympus Omniscan MX2
with a 2.25L16-AWS1 PAUT probe, which is a 2.25 MHz probe
with 16 active elements (active aperture of 16 mm by 16 mm).
Table 36 shows the results from this testing for all intended
flaws located in the test plates. This table includes the flaw
type, intended flaw height and length, the rejection rate from
the round robin results for conventional UT, the maximum
measured flaw length from the line scanning, and the maximum amplitude from the follow-up raster scanning. The flaw
length was measured using the 6dB drop method, and the flaw
length of scattered indications not separated by more than
2L was combined to determine the overall flaw length. Caution should be taken if comparing the results of this testing
to the current Annex K acceptance criteria as the reported
amplitude for each flaw was peaked during raster scanning.
Table 36. Experimental verification testing of flaw rejection amplitude limit.
Drawing Details
Flaw
Type
Flaw
Height
Flaw
Length
HAZ
Crack
0.18
0.52
UT
Rejection
Rate
NA
Raster Scan Results
Line Scan
Results
Rejection Limit
Maximum
Measured
Length
Maximum
Amplitude
-13 dB
-13 dB &
1" long
-10 dB
-8 dB
-6 dB
0.99
8.0
Y
N
Y
Y
Y
Porosity
0.10
0.73
NA
0.79
-6.2
Y
N
Y
Y
N
LOF
0.11
0.63
NA
1.14
10.8
Y
Y
Y
Y
Y
HAZ
Crack
0.14
0.57
NA
0.83
-0.6
Y
N
Y
Y
Y
Slag
0.10
0.74
NA
0.79
2.3
Y
N
Y
Y
Y
Crack
0.14
0.37
NA
0.51
-12.1
Y
N
N
N
N
LOF
0.18
0.50
NA
1.10
11.8
Y
Y
Y
Y
Y
LOF
0.12
0.64
NA
1.14
15.5
Y
Y
Y
Y
Y
0.19
0.49
NA
0.55
2.9
Y
N
Y
Y
Y
0.11
0.61
NA
0.87
-1.0
Y
N
Y
Y
Y
Slag
0.09
0.92
NA
0.63
-4.2
Y
N
Y
Y
Y
HAZ
Crack
0.14
0.45
NA
0.87
3.7
Y
N
Y
Y
Y
IP
0.10
0.88
NA
0.83
5.5
Y
N
Y
Y
Y
Slag
0.32
0.37
100%
0.55
1.7
Y
N
Y
Y
Y
Slag
0.16
0.18
100%
0.47
2.9
Y
N
Y
Y
Y
Toe
Crack
Toe
Crack
Slag
0.10
0.90
100%
1.73
-1.8
Y
Y
Y
Y
Y
Porosity
0.09
3.31
80%
3.39
-8.0
Y
Y
Y
Y
N
Toe
Crack
0.02
0.04
40%
0.28
-8.4
Y
N
Y
N
N
LOF
0.03
0.06
80%
0.36
-3.6
Y
N
Y
Y
Y
Slag
0.06
0.03
0%
0.39
-8.2
Y
N
Y
N
N
Slag
0.17
3.61
100%
3.47
0.9
Y
Y
Y
Y
Y
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
52
The maximum amplitude and measured length were used to
evaluate each flaw using five different criteria based on the
CIVA results:
1. Rejection of flaws with maximum amplitude ≥ –13 dB
2. Rejection of flaws with maximum amplitude ≥ –13 dB and
≥1″ long (i.e., reject scattered low-amplitude flaws such as
porosity)
3. Rejection of flaws with maximum amplitude ≥ –10 dB
4. Rejection of flaws with maximum amplitude ≥ –8 dB
5. Rejection of flaws with maximum amplitude ≥ –6 dB
All of the intended flaws had a maximum amplitude equal
to or greater than –13 dB. Therefore, all of the intended flaws
are rejectable under the first criteria (≥ –13 dB).
The second criteria was meant to reject low-amplitude
flaws, such as slag or porosity, that were over 1″ long and
would likely be applied in conjunction with another criteria
based solely on the amplitude. Six flaws are rejectable under
this criteria including three lack-of-fusion flaws, two slag
flaws, and one porosity grouping. All of the rejectable flaws
under this criteria were also rejectable to all of the other
amplitude-only criteria with the exception that the porosity would have been acceptable if the rejection limit was set
to –6 dB. Therefore, it may not be necessary to apply the
second criteria in conjunction with the other amplitudeonly criteria.
The third criteria (≥ –10 dB) rejected all intended flaws
except for a 0.14″ × 0.37″ crack. This crack is vertical and
embedded so the amplitude response relied on tip diffraction.
This specimen included weld reinforcement on both faces of
the weld, so the entire incidence angle range could not be
swept over the flaw before the front of the probe contacted
the weld reinforcement. Because of this, the incidence angles
were confined to high angles, and the maximum amplitude
was measured at the 67° incidence angle. CIVA analysis for
similar flaws had a maximum amplitude at ∼60° incidence
angle (since the reinforcement was not modeled in CIVA).
According to the critical flaw size for embedded planar flaws,
the critical flaw size for a similar aspect ratio for the 4 ksi stress
range in an equal thickness weld was approximately 0.13″ ×
0.31″. Therefore, this crack would have been critical, but since
this plate was not included in the round robin testing program,
the rejection rate according to conventional UT is unknown.
Three of the intended flaws are acceptable according to the
fourth criteria (≥ –8 dB). This includes the previously mentioned vertical, embedded crack along with a very small
surface breaking crack (0.02″ × 0.04″) and a very small slag
inclusion (0.06″ × 0.03″). The small surface breaking crack
had a rejection rate of 40% according to the conventional UT
round robin results while the small slag inclusion had a 0%
rejection rate for conventional UT.
Five of the intended flaws are acceptable according to the
fifth criteria (≥ –6 dB). This includes the three flaws from
the fourth criteria along with two groupings of porosity.
Once again, it should be mentioned that this testing was
performed on the round robin plates with the lowest amplitude indications. Therefore, it is likely that the indications
in the other round robin plates would have been rejected
with this criteria.
Finally, any amplitude-based acceptance criteria will have
variability from differences in probe parameters, calibration
procedures and standards, probe pressure, and final probe
location at maximum amplitude. Therefore, while verification testing is important to provide physical test results
to verify the CIVA modeling, specific maximum amplitude
values measured by a technician could be expected to vary
by ∼±4 dB.
3.7.4 Recommendation for AWS
A marked-up version of Annex K and associated commentary with the recommended changes is included in
Appendix G. Based on the CIVA analysis of critical planar
flaws and the experimental testing, it is recommended that
the acceptance criteria amplitude limit for flaw rejection be
set at 10 dB under the amplitude from the 1.5 mm (0.06″)
diameter SDH (i.e., –10 dB). As shown in Table 36, this limit
would result in rejection of all intended flaws from the verification testing except for the 0.14″ × 0.37″ embedded crack.
As stated previously, it is believed that the weld reinforcement that limited access—and thus limited the use of the
incidence angles that could be swept over this flaw—resulted
in the low amplitude response. Therefore, it is recommended
to set the Automatic Rejection Level (ARL) as 10 dB under
Standard Sensitivity Level (SSL) for tension welds. These
indications would be considered Class A defects and be
automatically rejected regardless of length.
The CIVA analysis of critical volumetric flaws found that
the amplitude associated with critical pores may be as low as
–13 dB. Volumetric flaws such as slag and porosity are typically
made up of scattered groupings of individual discontinuities.
In order to reject large groups of volumetric discontinuities
that include a critical sized pore, it is recommended that
indications with a maximum amplitude between –13 dB and
–10 dB during the follow-up raster scanning be rejected
if the length of the entire grouping of discontinuities
exceeds 1″. It is recommended that a new amplitude limit
referred to as the Evaluation Level (EVL) be set at 13 dB under
SSL, and indications which exceed the EVL but are less than
the ARL be defined as Class B indications.
Measurement of the flaw length will be needed for rejection evaluation of Class B indications or limits for repair
of Class A defects. It is recommended that the length
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
53 measurement for flaws use the 6 dB drop method during
the manual raster scan. Some PAUT acceptance criteria use
a standard amplitude limit for length measurement rather
than the 6 dB drop. In these cases, the length is determined
to encompass the full extent of the flaw, which has amplitude greater than this limit. While use of this method has
some merit, oversizing of indications with saturated signals
may occur.
It is worth commenting that the research team initially
intended on recommending that an amplitude limit of
18 dB below SSL be used for the length measurement of indications which are above the EVL but less than the ARL (i.e.,
Class B). (The –18 dB limit corresponds to the amplitude limit
for detection of flaws during the encoded line scan, as will be
described in the following section.) In this case, length measurement of Class B flaws would have involved measuring
the extents where the signal exceeds the flaw detection limit.
In essence, this would have ensured that any indications that
would be considered as detectable are included in the evaluation against the 1″ length limit. Since 6 dB drop on flaws
with maximum amplitude between –13 dB and –10 dB will
essentially be equivalent to the –18 dB limit (i.e., –19 dB and
–16 dB, respectively), it was finally decided to recommend the
use of the familiar 6 dB drop method for all length measurements. This will provide consistency for Class A and Class B
flaws, along with flaws in compression welds.
It should also be noted that the existing requirements on
spacing between Class B flaws from each other and the edge
of the plate have been retained.
Finally, it is recommended that for all indications investigated in follow-up manual raster scanning, a screenshot
should be required and data documented at the location of
maximum indication amplitude.
3.8 Amplitude Limit for
Detection of Flaws
3.8.1 CIVA Modeling
As stated previously, the approach to the inspection procedure is to use line scanning for detection of weld flaws
and raster scanning to evaluate rejection of weld flaws. The
amplitude limit for detection of weld flaws will need to
be set such that it will detect critical weld flaws during the
line scan regardless of the flaw tilt, skew, and position in relation to the probe. The detection amplitude limit, by definition, must be lower than the acceptance criteria rejection
limit since the rejection limit is compared to the maximum
possible amplitude following raster scanning.
In order to develop the amplitude limit for flaw detection,
CIVA analysis was performed using the same probe that was
used in the flaw rejection limit (2.25 MHz PAUT probe with
an active aperture of 16 mm by 16 mm and an angular range
of 45°–70°). The CIVA analysis was performed on the same
size flaws used in the flaw rejection, which are similar in size
to the critical planar and volumetric flaws. Since the amplitude was always referenced to the 1.5 mm (0.06″) diameter
SDH at a similar sound path, the reported amplitude is similar to that which would be found during typical line scanning
with TCG. Positive amplitude represents a higher amplitude
response than the SDH, and negative amplitude represents a
smaller amplitude response than the SDH.
Encoded line scanning involves using the incidence angle
range of PAUT focal laws to provide coverage of the entire
weld volume and HAZ. As shown in Figure 23, the PAUT
probe is moved in a direction parallel to the weld axis at a
constant index offset and with the probe orientation remaining perpendicular to the weld axis.
Due to the amplitude being strongly affected by the
interaction of the flaw tilt and sound beam incidence angle,
full coverage of the weld volume and HAZ should be provided in two crossing directions (i.e., nearly perpendicular
sound beam directions), as shown in Figure 24. This can
be provided by either scanning from both sides of the weld
or combining 1st and 2nd leg index offset scans from the
same side of the weld. Due to the effects of attenuation and
beam spread on long sound paths, scanning from both sides
of the weld is preferred rather than relying on second leg
scans with long sound paths. This has already been incorporated in Annex K by requiring butt welds be tested from
the same face but opposite sides of the weld axis where
access is possible.
JIS Z 3060 [13], which uses a DAC curve approach, requires
that the plate be flipped and scanning performed in the 1st leg
from the other face of the weld when the sound path exceeds
250 mm (9.8″). For a 70° incidence angle, this would correspond to a depth of 3.4″ which would be exceeded for
2nd leg scanning of plates thicker than 1.7″. Therefore, it may
be reasonable to set limits on the maximum sound path which
can be used for sound coverage in order to limit the effects
of attenuation and beam spread. For instance, coverage could
still be provided for shorter sound paths through the use of
additional line scans at a different index offset or flipping the
Figure 23. Encoded line scanning.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
54
Figure 24. Line scanning sound coverage.
plate and scanning from both weld faces. A reasonable limit
may be limiting the sound path used for full coverage to 12″,
since this would still allow for full coverage to be provided
at the 70° incidence angles for the 2nd leg in 2″ thick plates
(i.e., 4″ deep TCG point). The recommendations for changes
to Annex K included a statement that the probe dimensions
shall be chosen in order to optimize the beam formation
within the area of coverage. No exact sound path limit was
provided.
In order to develop the amplitude limit for the detection of
flaws, the probe was moved across the weld flaws perpendicular to the weld axis (i.e., the index offset was varied) in the
CIVA simulations. The largest amplitude for all of the focal
laws (incidence angles) was documented at 6 mm (0.24″)
increments of the index offset, as shown in Figure 25. The
maximum amplitude across the incidence angle range for each
index offset represents the largest amplitude that would occur
if a line scan was performed using the same index offset. The
indication amplitude was documented at a small increment
of possible index offsets. The amplitude limit for flaw detection was subsequently determined in order to detect critical
weld flaws for any possible index offset used in line scanning
(i.e., combination of possible probe and flaw locations). The
only stipulation is that the flaw must be within the coverage of the incidence angle range, which is already provided
through minimum scanning coverage requirements. In other
words, as long as full coverage was provided of the weld and
HAZ, the amplitude from the critical weld flaw would surpass
this limit and therefore be detected during line scanning.
Since the flaw tilt is also an unknown parameter, the tilt of
planar flaws was evaluated at 0°, ±5°, ±30°, and ±45°. Flaw
tilt away from the probe (the maximum amplitude in 1st leg)
was defined as positive tilt while tilt towards the probe (the
maximum amplitude in 2nd leg) was defined as negative tilt,
as shown in Figure 26. Since the minimum scanning coverage requirements provide for full coverage in two crossing
directions, flaws tilted away from the probe when scanned
from one side of the weld would be tilted towards the probe
when scanned from the other side of the weld. Therefore, all
necessary analysis could be performed by sweeping the probe
over the flaw in 1st and 2nd leg. The smaller of the amplitude
from the 1st leg results for positive flaw tilt or 2nd leg results
for negative flaw tilt was used as the controlling amplitude
for flaw detection in order to ensure that the flaw would be
detected by only requiring sound coverage in two crossing
directions. Therefore, it did not matter whether the sound
beam that would impact the flaw was provided in 1st or 2nd
leg as long as the full weld volume and HAZ was covered by
sound in two crossing directions.
Figure 27 displays the amplitude of a 0.2″ × 0.2″ embedded
planar flaw at mid-thickness depth in a 2″ thick plate, as the
probe is swept over the flaw in 1st and 2nd leg. When there is
no tilt, the amplitude varies from –5 dB to –15 dB compared
to the 1.5 mm (0.06″) diameter SDH until the weld flaw is no
Figure 25. Index offset increment.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
55 Figure 26. Flaw tilt during line scanning.
longer within the sound beam coverage, at which point the
amplitude drops quickly (0″ index offset). When the flaw is
tilted from vertical by 30° or 45°, the maximum amplitude
increases to +10 dB when the sound beam is nearly perpendicular to the weld flaw. For the negative tilt cases, this occurs
at index offsets of –4″ to –2″ where the flaw is impacted by
sound in the 2nd leg. For positive tilt cases, it occurs at index
offsets of –1″ to 0″ where the flaw is impacted by sound in
the 1st leg. When the sound beam is nearly parallel to the
tilted flaws (1st leg of negative tilted flaws and 2nd leg of positive tilted flaws), the amplitude drops off considerably with
amplitudes of ∼–25 dB. Therefore, sound coverage in two
crossing directions is required in order to provide for detection of tilted flaws.
The critical flaws from Table 33 and Table 34 for the CIVA
analysis used to develop the acceptance criteria rejection
limits were also used in the study for the amplitude limits for
flaw detection. The controlling amplitude over all possible
index offset positions was tabulated for various flaw tilt and
possible incidence angle ranges. The controlling amplitude
was used to determine the overall minimum amplitude possible during line scanning for critical weld flaws. This amplitude
could then be used to form the basis of the amplitude detection limit for encoded line scanning, which would require
follow-up raster scanning for evaluation of acceptance.
Table 37 includes the results from the CIVA flaw detection
analysis for planar flaws without any skew (i.e., the flaw length
is parallel to the weld axis). This table presents the minimum
possible peak amplitude during a line scan as long as the flaw
was within sound coverage provided in two crossing directions. It presents parametric results for various-sized planar
flaws, various flaw tilt, and various incidence angle ranges. The
controlling line scan amplitudes (i.e., minimum peak amplitude depending on the chosen index offset) are highlighted for
each flaw and incidence angle range. The maximum amplitude (from raster scanning) is also provided in the table.
For some of the flaws, especially those with large tilt, the
amplitude is very sensitive to the probe location. For instance,
0.20" x 0.20" Tilted Embedded Flaw in 2" Plate
Change from Reference Amplitude (dB)
15
10
5
0
0.20"x0.20" 0 Tilt
-5
0.20"x0.20" 5 Tilt
-10
0.20"x0.20" -5 Tilt
-15
0.20"x0.20" 30 Tilt
-20
0.20"x0.20" -30 Tilt
-25
-30
0.20"x0.20" 45 Tilt
-35
0.20"x0.20"x -45 Tilt
-40
-7
-6
-5
-4
-3
-2
Index Offset (in)
-1
0
1
Figure 27. Amplitude profile as probe is swept over weld flaw of various tilt angles.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
56
Table 37. CIVA flaw detection results for planar flaws.
Flaw Parameters
Location
Height
(in)
0.20
0.15
Length
(in)
0.20
0.15
Plate
Thickness
(in)
2
2
Embedded
0.10
0.10
0.10
0.10
0.10
0.15
0.5
2
0.5
Surface
0.05
0.10
0.5
Tilt
(deg)
Maximum
Amplitude
(dB)
Minimum Peak Line Scan Amplitude
for Incidence Angle Range (dB)
45° 70°
50° 70°
55° 65°
45° 55°
45° 50°
0
-5
-16
-16
-14
-15
-15
+/- 5
-6
-16
-12
-12
-16
-16
+/- 30
9
-10
-10
-9
-9
-9
+/- 45
9
-12
-12
-10
-6
4
0
-7
-17
-15
-15
-17
-17
+/- 5
-8
-17
-13
-13
-17
-17
+/- 30
5
-11
-10
-1
-11
-11
+/- 45
5
-15
-15
-12
-8
0
0
-6
-14
-14
-14
-10
-10
+/- 5
-10
-13
-13
-12
-13
-13
+/- 30
-4
-11
-11
-8
-11
NA
+/- 45
-6
-15
-15
-15
-13
-6
0
-10
-21
-20
-17
-21
-21
+/- 5
-10
-20
-18
-17
-20
-20
+/- 30
-2
-14
-10
-7
-14
-14
+/- 45
-2
-17
-17
-17
-8
-5
0
3
-3
-3
1
-3
-3
+/- 5
-2
-8
-8
-8
-5
-4
+/- 30
0
-13
-13
-7
-10
-10
+/- 45
-2
-17
-17
-15
-6
-6
0
-3
-11
-9
-6
-11
-11
+/- 5
-6
-12
-12
-12
-11
-11
+/- 30
-8
-15
-15
-11
-15
-9
+/- 45
-10
-15
-15
-15
-11
-11
the 0.20″ × 0.20″ embedded flaw with 45° tilt has a maximum
amplitude of +9 dB from raster scanning, but a minimum peak
amplitude during possible line scan locations of –12 dB even
with full coverage from two crossing directions. Therefore, the
amplitude of this flaw could be 21 dB below the maximum
during the line scanning, even with providing full coverage in
two crossing directions. This highlights the need for follow-up
raster scanning rather than evaluating flaw rejection on the
line scan results.
The incidence angle ranges shown in Table 37 were chosen
to investigate whether limits should be placed on the incidence angle used in the scan plan. Limiting the incidence
angle range, while providing less sound coverage, may result
in larger amplitudes for flaw detection. This would be similar to the requirement in Annex K where the incidence angle
used in the line scan must be within ±10° of the weld fusion
face. Rather than perform five different analyses for each flaw
(one for each incidence angle range investigated), the individual incidence angle that had the largest amplitude across
the 45°–70° angular range was documented for each index
offset evaluated (0.24″ increment). The minimum amplitude
for each incidence angle range could then be determined
from this data since, as the probe is swept over the flaw, the
maximum amplitude at each index point will occur at a
slightly different incidence angle.
For example, referring back to the 0.20″ × 0.20″ flaw with
45° tilt, the maximum amplitude is +9 dB, but the minimum
amplitude during line scanning using the 45°–70° incidence
angle range was –12 dB. It is anticipated that the maximum
amplitude would occur at an incidence angle of 45° since this
would be perpendicular to the flaw. Therefore, limiting the
incidence angle range closer during the line scan to 45° should
result in a larger amplitude. This was confirmed in the CIVA
analysis, where the minimum amplitude over the 45°–55°
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
57 Table 38. CIVA flaw detection results for volumetric flaws.
Flaw Parameters
Controlling
Maximum
Plate
Amplitude
Height Length
Thickness
Location
(dB)
(in)
(in)
(in)
0.25
0.25
0.5
-7
0.25
0.25
2
-6
0.125
0.125
0.5
-13
Embedded
0.125
0.125
2
-13
0.08
0.08
0.5
-18
0.08
0.08
2
-18
0.25
0.25
0.5
-7
0.25
0.25
2
-5
0.08
0.08
0.5
-13
Surface
0.08
0.08
2
-11
0.03
0.03
0.5
-18
Minimum Line Scan Amplitude for
Incidence Angle Range (dB)
45°–
70°
50°–
70°
55°–
65°
45°–
55°
45°–
50°
-13
-13
-16
-20
-21
-25
-12
-12
-18
-22
-28
-13
-13
-16
-20
-21
-25
-9
-12
-18
-22
-28
-10
-12
-16
-18
-21
-23
-9
-8
-17
-18
-22
-13
-10
-15
-17
-21
-22
-12
-8
-15
-15
-18
-13
-10
-15
-17
-21
-21
-12
-6
-15
-13
-18
index offset). For instance, the minimum amplitude during
line scanning is relatively unaffected by the different combinations of incidence angle ranges evaluated. Still, a loss of
amplitude of 7 dB was typical for the line scan amplitude
compared to the maximum possible during raster scanning.
Similar to the raster scanning amplitude limits, the CIVA
results could be further summarized by combining the results
for flaw sizes comparable to the critical flaw sizes. The flaw
detection limits for the 4 ksi and 8 ksi stress ranges for equal
thickness and thickness transition welds is shown in Table 39.
For planar flaws, limiting the incidence angle range to 55°–65°
resulted in the largest line scan amplitudes, but only by a
few decibels compared to using 45°–70°. Using an incidence
angle range of 55°–65° would result in much less coverage than using 45°–70°. Therefore, based on these results, it
seems that using an incidence angle range from 45°–70° is justified without the need for additional scan plan requirements
beyond providing full coverage of the entire weld volume and
HAZ in two crossing directions.
As explained previously and shown in Figure 23, encoded
line scanning is performed by keeping the probe perpendicular to the weld axis while probe movement is parallel
to the weld axis. All of the previous CIVA analysis assumed
incidence angular range was –6 dB and over the 45°–50° incidence angular range was +4 dB. Obviously, the amplitude of
this flaw is very sensitive to the incidence angle, as is typical
for tilted lack-of-fusion flaws.
Along with limiting the sound coverage, limiting the incidence angle range assumes that the flaw tilt is known. Lack
of fusion flaws are typically assumed to have the same tilt as
the fusion face. However, one must also consider if the flaw is
a vertical crack or tilted at ±5°. Returning once again to the
example of the 0.20″ × 0.20″ flaw, if the weld had a 45° bevel
face and the incidence angular range was limited to either
45°–55° or 45°–50° but the weld flaw was tilted at ±5°, the
minimum peak line scan amplitude would be –16 dB for all
possible index offsets. Therefore, using a flaw detection limit
of –6 dB or +4 dB would be unconservative and the critical
weld flaw would not be detected. While limiting the incidence
angle range may be helpful for flaws with known tilt in order
to better maximize the amplitude, it does not help when the
flaw tilt is vertical or otherwise unknown.
Table 38 includes the results from the CIVA flaw detection
analysis for volumetric flaws. Since these flaws were modeled
as spherical porosity, there is no flaw tilt or skew. As expected,
spherical flaws are not as sensitive to probe location (i.e.,
Table 39. Summary of CIVA flaw detection amplitude results without
flaw skew.
4 ksi (Equal)
4 ksi (Transition)
8 ksi (Equal)
8 ksi (Transition–
Embedded)
Planar Flaw Amplitude (dB)
45°– 50°– 55°– 45°– 45°–
70°
70°
65°
55°
50°
-17
-17
-14
-16
-16
-17
-16
-15
-17
-17
-17
-16
-15
-17
-17
Volumetric Flaw Amplitude (dB)
45°– 50°– 55°– 45°– 45°–
70°
70°
65°
55°
50°
-13
-13
-12
-13
-13
-22
-22
-18
-15
-15
-22
-22
-18
-17
-17
-21
-25
-20
-17
-21
-21
-25
-23
-22
Copyright National Academy of Sciences. All rights reserved.
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
58
Figure 28. Flaw skew.
that the flaw was aligned parallel to the weld axis and, therefore, the probe was perfectly perpendicular to the weld flaw.
While this is a valid assumption for raster scanning where
the probe will be rotated as well as translated, it may be
unconservative for line scanning.
In order to account for this effect, the change in amplitude
due to flaw skew was evaluated. CIVA analysis of embedded and
surface vertical weld flaws in 0.5″ and 2″ plates with 5°, 10°,
and 20° skew was performed and compared to the results with
no skew. Since the sound is reflected to the side of skewed planar flaws, as shown in Figure 28, lateral movement of the probe
along the weld axis was performed as well as sweeping the probe
over the flaw perpendicular to the weld axis. (Note: this figure
represents the centerline of the sound beam, but the beam actually has beam spread and width.)
The maximum amplitude for flaws with skew is shown in
Figure 29 for surface flaws and Figure 30 for embedded flaws.
As the flaw skew is increased, the drop in amplitude increases,
especially for larger flaws. This is due to the fact that the beam
is hitting different parts of the flaw at different times, which
causes the amplitude to drop more severely. Flaws with long
sound paths also had larger drops in amplitude since the
beam reflected off of the skewed flaw travels a further distance
transverse to the probe. For instance, the embedded flaws
in the 2″ thick plate had greater drop in amplitude than the
embedded or surface flaws in the 0.5″ thick plate. The flaw
and plate combinations chosen for this analysis correspond
to those included in Table 37.
In order to account for the effect of flaw skew on flaw
detection, the drop in amplitude from the analysis of the vertical flaw with skew was added to the results of the tilted flaws
without skew for each individual flaw. This assumes that the
drop in amplitude from flaw skew will be similar for tilted
and vertical flaws.
This assumption was checked for the 0.20″ × 0.20″ embedded flaw in a 2″ thick plate by modeling flaws with both skew
and tilt and comparing the results to the estimated values.
It was determined that this assumption was reasonable. For
instance, in the case of 5° skew, the drop in amplitude due to
skew on the vertical flaws was –4 dB. The drop in amplitude
due to skew on the tilted flaws varied from –2 dB to –5 dB
depending on the flaw tilt. In the case of a 10° skew where
the drop in amplitude due to skew on the vertical flaws was
–12 dB, the drop in amplitude for the tilted flaws varied from
–7 dB to –15 dB.
The flaw detection limits, including the effect of skew on
the planar flaws, are given in Table 40 for 5° skew and Table 41
Change from Reference Amplitude (dB)
10
5
0
-5
0.05"x0.05" Surface Flaw in 0.5" Plate
-10
0.05"x0.10" Surface Flaw in 0.5" Plate
-15
0.05"x0.15" Surface Flaw in 0.5" Plate
-20
0.10"x0.10" Surface Flaw in 0.5" Plate
-25
0.10"x0.15" Surface Flaw in 0.5" Plate
-30
-35
0
5
10
15
Flaw Skew (deg)
20
25
Figure 29. Flaw skew results for surface flaws.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
59 Change from Reference Amplitude (dB)
10
5
0
-5
0.10"x0.10" Embedded Flaw in 2" Plate
-10
0.10"x0.30" Embedded Flaw in 2" Plate
-15
0.20"x0.20" Embedded Flaw in 2" Plate
-20
0.15"x0.15" Embedded Flaw in 2" Plate
-25
0.10"x0.10" Embedded Flaw in 0.5" Plate
-30
-35
0
5
10
15
Flaw Skew (deg)
20
25
Figure 30. Flaw skew results for embedded flaws.
Table 40. Summary of CIVA flaw detection amplitude results
with 5ç flaw skew.
4 ksi (Equal)
4 ksi (Transition)
8 ksi (Equal)
8 ksi (Transition—
Embedded)
Planar Flaw Amplitude (dB)
45°– 50°– 55°– 45°– 45°–
70°
70°
65°
55°
50°
-20
-20
-18
-20
-20
-19
-17
-17
-19
-19
-19
-17
-17
-19
-19
Volumetric Flaw Amplitude (dB)
45°– 50°– 55°– 45°– 45°–
70°
70°
65°
55°
50°
-13
-13
-13
-12
-13
-15
-22
-22
-18
-15
-17
-22
-22
-18
-17
-23
-25
-22
-19
-23
-23
-25
-23
-22
-21
Table 41. Summary of CIVA flaw detection amplitude results
with 10ç flaw skew.
4 ksi (Equal)
4 ksi (Transition)
8 ksi (Equal)
8 ksi (Transition—
Embedded)
Planar Flaw Amplitude (dB)
45°– 50°– 55°– 45°– 45°–
70°
70°
65°
55°
50°
-28
-28
-26
-28
-28
-26
-24
-24
-26
-26
-26
-24
-24
-26
-26
Volumetric Flaw Amplitude (dB)
45°– 50°– 55°– 45°– 45°–
70°
70°
65°
55°
50°
-13
-13
-13
-12
-13
-15
-22
-22
-18
-15
-17
-22
-22
-18
-17
-25
-25
-24
-22
-25
-25
-25
-23
-22
Copyright National Academy of Sciences. All rights reserved.
-21
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
60
for 10° skew. In general, including the 5° skew resulted in a
–2 dB to –3 dB decrease in the flaw detection amplitude while
the decrease was –9 dB to –12 dB for 10° skew. Since the drop
in amplitude due to skew is greater for larger flaws, the 4 ksi
equal thickness weld had lower flaw detection amplitude than
the 4 ksi transition or 8 ksi equal thickness welds. Since the
volumetric flaws were assumed to be spherical, the volumetric
flaws were not affected by flaw tilt or skew; therefore, the
volumetric flaw results are the same as those in Table 39.
The proposed PAUT annex for AWS D1.1 was reviewed
by the research team after performing the CIVA analysis.
The proposed PAUT annex for D1.1 utilizes an incidence
angle range from 40°–70° rather than the 45°–70° incidence angle range in D1.5 Annex K. One of the authors of
the proposed D1.1 PAUT annex was contacted in order to
obtain an explanation for the increase in the incidence angle
range from 45° to 40°. The purpose of the incidence angle
range increase was to enlarge the sound coverage area and to
aid in verification of corner trap signals. The AWS D1.1 proposed PAUT annex also includes a requirement that the HAZ
be covered with incidence angle range from 40°–60° in order
to increase the detectability of corner trap signals from surface
breaking HAZ cracks. The author of the proposed D1.1 PAUT
annex stated that incidence angle was limited at 40° since
standing wave signals were produced sometimes at 35° and
this seemed risky since it was close to the first critical angle.
Subsequently, additional CIVA analysis was performed by
the research team on a subset of flaws with an extended incidence angle range of 40°–70°. This was used to verify that the
previously determined flaw detection amplitude limits using
the 45°–70° incidence angle range would be appropriate for
use with a 40°–70° incidence angle range. All of these results
with the 40°–70° range were within ±1 dB of the results given
in Table 39 for flaws without skew. Therefore, it was determined that the incidence angle range could be extended from
45°–70° to 40°–70° without significantly altering the necessary flaw detection limits.
3.8.2 Verification Testing of Weld
Flaw Samples
In order to verify the CIVA results, physical testing was performed on weld samples with known weld flaws. These specimens included the round robin test plates as well as additional
test plates with small weld flaws. None of the plates were fabricated with acoustically anisotropic material. This testing was
performed in conjunction with final verification testing of the
flaw rejection limits by an ASNT Level III UT/Level II PAUT
technician. The weld samples were scanned with an Olympus
Omniscan MX2 with a 2.25L16-AWS1 PAUT probe, which is
a 2.25 MHz probe with 16 active elements (active aperture of
16 mm × 16 mm). Some supplemental line scanning was per-
formed with a 5L64-A12 PAUT probe, which is a 5 MHz probe
with 32 active elements (active aperture of 19.2 mm × 10 mm)
in order to evaluate the differences between the 2.25 MHz and
5 MHz probes.
The weld samples were scanned after performing TCG
calibration on an AISI 1018 calibration block with 1.5 mm
(0.06″) diameter SDHs. The reference amplitude was set
to 80% FSH, and +12 dB scanning gain was added. Line
scanning was performed using an incidence angle range of
40°–70° with full coverage of the weld volume and HAZ in two
crossing directions. No additional requirements were imposed
on the scan plan. All indications with an amplitude greater than
–20 dB were further investigated through raster scanning for
flaw rejection verification testing, as previously explained.
Since some of the plates had reinforcement on both faces of
the plate, these plates were line scanned in the 1st leg from each
side of the weld and each face of the plate (i.e., four line scans).
This was also performed for the plates which were thicker than
0.75″, since the TCG did not extend beyond 2″ depth. Subsequently, the 2nd leg portion of the scan would have extended
beyond the last TCG point. Scanning of the 0.75″ thick plates
was performed with two line scans: one from each side of the
weld using the 1st and 2nd leg to cover the entire weld, and
HAZ in two crossing directions. Some of these plates had
additional scans performed from the other face of the plate to
verify whether all of the intended weld flaws would still have
been detected if full coverage was provided from that face.
The results of the verification testing are shown in Table 42
for the intended weld flaws. This table includes the maximum
amplitude for each flaw after evaluating each of the line scans
necessary for full coverage in two crossing directions. It also
includes the maximum amplitude from the follow-up raster
scanning that was previously reported. All of the results provided in the table are from testing with the 2.25 MHz probe,
but the amplitude with the supplemental line scans with the
5 MHz probe was found to be similar. Since +12 dB scanning
gain was added to the line scans and the Olympus Omniscan
MX2 instrument truncates the A-scan at 250% FSH, the maximum amplitude during the line scan was truncated at –2.1 dB.
As expected, the maximum amplitude from the line scan
can be significantly lower than the maximum amplitude
from the raster scanning. The largest difference between the
line scan and the raster scan (excluding truncated line scan
results) was for the slag, which was 0.16″ × 0.18″. This flaw
had a maximum amplitude during the line scans from Face
A of the plate of –9.9 dB, while the follow-up raster scan was
+2.9 dB. This is a difference of 12.8 dB, approximately a factor
of four times as much amplitude.
The intended flaw with the smallest maximum amplitude
after evaluating each line scan necessary for full coverage was
–13.9 dB for the embedded vertical crack, which was 0.14″ ×
0.37″. Therefore, all of the intended flaws would be detected
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
61 Table 42. Verification testing of flaw detection results.
Drawing Details
Flaw
Flaw
Flaw Type
Height (in) Length (in)
HAZ Crack
0.18
0.52
Porosity
0.10
0.73
LOF
0.11
0.63
HAZ Crack
0.14
0.57
Slag
0.10
0.74
Crack
0.14
0.37
LOF
0.18
0.50
LOF
0.12
0.64
Toe Crack
0.19
0.49
Toe Crack
0.11
0.61
Slag
0.09
0.92
HAZ Crack
0.14
0.45
IP
0.10
0.88
Slag
0.32
0.37
Slag
0.16
0.18
Slag
0.10
0.90
Porosity
0.09
3.31
Toe Crack
LOF
Slag
Slag
0.02
0.03
0.06
0.17
0.04
0.06
0.03
3.61
Line Scan
Results
Maximum
Amplitude (dB)
-2.1*
-10.5
-2.1*
-2.1*
-6.4
-13.9
-2.1*
-2.1*
-2.1*
-2.1*
-3.0
-2.1*
-2.1*
-2.1* (Face A)
-2.1* (Face B)
-9.9 (Face A)
-2.1* (Face B)
-3.4 (Face A)
-2.1* (Face B)
-8.7 (Face A)
-12.5 (Face B)
-8.3
-2.8
-12.1
-2.1*
Raster Scan
Results
Maximum
Amplitude (dB)
8.0
-6.2
10.8
-0.6
2.3
-12.1
11.8
15.5
2.9
-1.0
-4.2
3.7
5.5
1.7
2.9
-1.8
-8.0
-8.4
-3.6
-8.2
0.9
*A-scan was truncated at 250% FSH with +12 dB scanning gain, which correlates to -2.1 dB
for any flaw detection limit of –14 dB or less. All of the flaw
detection limits for planar flaws determined through the CIVA
analysis and presented in Table 39 are below –14 dB, so use of
the CIVA results would have resulted in all of the intended
flaws to be detected. None of the weld flaws in the weld flaw
samples had any skew relative to the weld axis. For this reason,
the maximum line scan amplitude should be compared to the
limits in Table 39 rather than those in Table 40 or Table 41.
During the verification testing, there were many unintended
indications that crossed the initial detection limit threshold of
–20 dB, excluding geometric indications from weld reinforcement and surface roughness. While some of these unintended
flaw detections seem to correlate to actual unintended flaws,
most of these indications seemed to be spurious repeating signals on the high incidence angle (70°), as shown in Figure 31
(left). Sometimes these repeating signals also appeared at the
Figure 31. Spurious signals on high-incidence (left) and low-incidence (right) angles.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
62
low incidence angle (40°), as shown in Figure 31 (right). These
spurious indications also appeared during scanning of clean
production welds that had been inspected with digital RT and
had no noted indications. The spurious indications were only
noticed in the scanning with the 2.25L16-AWS1 probe and not
with the 5L64-A12 probe. After further investigation, it was
determined that these indications were likely due to grating
lobes due to the large pitch of the 2.25 MHz probe (1 mm)
compared with the wavelength (1.45 mm). From the literature
[47], grating lobes typically appear when the pitch is greater
than the wavelength, although they may appear for a slightly
smaller pitch as well. The spurious indications seemed sensitive
to changes to the surface roughness of the plates, even though
the surface roughness was typical of production welds. Grating
lobes may reflect off of the surface roughness, which may
explain why they appeared on locations with more roughness.
Follow-up raster scanning on the detected locations of the spurious indications found no relevant indications greater than
–20 dB, which confirms that they are noise. Due to the prevalence of these grating lobe signals throughout the line scans, it
was determined to exclude indications that were characterized
as grating lobe indications from follow-up raster scanning.
Table 43 summarizes the detection and rejection results of
unintended indications during the verification testing. When
the flaw detection limit was set at –20 dB, 36 unintended indications (excluding geometric indications) were detected that
required follow-up raster scanning. Of those 36 indications,
33 indications had a maximum amplitude during raster scanning that was less than –13 dB; these indications would be
accepted under all of the flaw rejection limits previously discussed. Three unintended indications were detected above
–20 dB and would be rejected if the flaw rejection limit were set
at –13 dB, while only one of these indications would be rejected
if the flaw rejection limit were set at –10 dB or higher.
As the flaw detection limit is shifted up (more positive),
the number of unintended indications drops dramatically,
but indications may be missed that would have been rejected
during follow-up raster scanning. For instance, if the flaw
detection limit were set at –18 dB, the number of acceptable unintended indications that are detected and require
follow-up raster scanning decreases from 33 to 18. Increasing the flaw detection limit to –16 dB decreases the number
of acceptable unintended indications to only 8, but one of
the unintended indications that would have been rejectable
if the rejection limit were set to –13 dB would not have been
detected. Thus, that indication would now be accepted since
it was effectively missed.
Due to the high rate of unintended indications found
during the testing of the weld flaw specimens, a clean production weld 32 inches long by 2 inches thick was obtained
and scanned. This specimen was specifically obtained to get
a feel for what level of “noise” might be expected in clean
shop production welds. Other than grating lobe signals, only
two indications (excluding geometric indications) were
noted as greater than –20 dB during four line scans with the
2.25 MHz probe, and no indications (excluding geometric
indications) appeared above this threshold with the 5 MHz
probe. Neither of the two indications detected with the
2.25 MHz probe exceeded the –18 dB detection limit.
Therefore, excluding the grating lobe indications, it is
anticipated that very few unintended indications would be
identified above a flaw detection threshold of –18 dB in
clean welds.
3.8.3 Recommendation for AWS
A marked-up version of Annex K and associated commentary with the recommended changes is included in Appendix G. Based on the CIVA results for planar flaws along with
the verification testing, it was determined to set the amplitude limit for flaw detection at 18 dB under the SSL (–18 dB).
In Annex K, this is referred to as the disregard level (DRL)
because indications with amplitudes lower than this limit
during the line scans will not require additional raster scanning. As stated, this limit provided adequate sensitivity for all
intended flaws in the verification testing, along with detecting
all unintended indications that were –13 dB or greater when
performing follow-up raster scanning. Therefore, this limit
seems to set a good compromise between adequate detection
of critical flaws and adequate sensitivity, so that the number
Table 43. Number of unintended indications detected
during verification testing.
Raster Scanning Amplitude
Threshold
Accepted (<-13 dB)
Rejected > -13 dB
Rejected > -13 dB & 1" long
Rejected > -10 dB
Rejected > -8 dB
Rejected > -6 dB
Line Scan Amplitude Thresholds
Detected
Detected
Detected
Detected
at -20 dB
at -18 dB
at -16 dB
at -14 dB
33
18
8
2
3
3
2
1
0
0
0
0
1
1
1
1
1
1
1
1
1
1
1
1
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
63 of harmless indications which require follow-up raster scanning is minimal.
According to the CIVA analysis, the –18 dB limit would
overestimate the lowest possible amplitude from a critical
flaw with skew. This may possibly result in missing a critical
flaw, but based on the verification testing, it seems that the
CIVA results for flaw detection were slightly conservative.
For instance, the CIVA results for a flaw without skew showed
that an amplitude limit of –17 dB would result in detection of
all critical flaws, but the minimum amplitude measured during
verification testing for a comparable flaw was actually –14 dB.
Assuming the amplitude of flaws with 5° skew are comparably
overestimated, a flaw detection limit of –18 dB should detect
critical flaws with 5° skew as the CIVA analysis for these flaws
gave an amplitude limit of –20 dB. In other words, no further
reduction below SSL seems to be required.
Flaws with 10° skew or greater may have much lower amplitude than the flaw detection limit proposed. For instance, the
CIVA result for 10° skew was –28 dB and would be much
lower for 20° skew. Setting an amplitude limit this low would
result in a large amount of indications which would be
acceptable and would likely result in manual raster scanning
of most, if not all, of the weld. This would eliminate the economic advantage of encoded line scanning. The likelihood
of planar flaws with skew is expected to be low since the LOF
and incomplete penetration flaws will likely be aligned along
a fusion face. Also, the current requirements for follow-up
scanning for transverse flaws using scanning Pattern D or E
will be retained. These requirements will allow for detection
of flaws transverse to the weld axis and may aid in detecting
highly skewed flaws.
It is recommended that 12 dB over SSL be added to the scanning gain during encoded line scanning of tension welds in
order to provide adequate screen height of indications greater
than the flaw detection limit. After applying 12 dB of scanning
gain, an indication which is greater than 40% FSH will require
follow-up raster scanning, assuming that reference amplitude
is set at 80% FSH.
3.9 Compression Weld
Acceptance Criteria
The flaw detection and rejection limits discussed previously
were derived using the critical flaw sizes. These sizes were computed assuming that the weld is in a tension member for the FFS
analysis. AWS traditionally has separate acceptance criteria for
tension and compression welds for both conventional UT and
Annex K. For conventional UT in accordance with Clause 6,
the rejection limits for compression welds are ∼6 dB lower (less
conservative) than for tension welds. There is also a slight modification to the Class C length limits, as the maximum length
for compression welds is 2″ regardless of the through-thickness
location, while there are tighter limits for tension welds in the
top or bottom quarter of the weld thickness. For Annex K, the
amplitude limits are exactly the same for compression and
tension welds, but the maximum length limits are slightly
different for compression and tension welds. The Class C length
limits are carried over from Clause 6 with the smaller length for
near surface flaws in tension welds. The Class B length limits
are also slightly different, with compression welds having a
maximum length of 0.75″ and tension welds having a maximum length of 0.5″.
Rather than include separate acceptance criteria for tension and compression welds, the CSA W59:2018 code only
includes separate acceptance criteria for statically loaded and
cyclically loaded structures. Therefore, while bridges would
fall under cyclically loaded, tension and compression welds
would be evaluated using the same acceptance criteria.
Since the critical flaw size of compression welds was not specifically determined (and could not be using FFS), the acceptance criteria for compression welds would either be based
on workmanship criteria or on the results for tension welds.
It is recommended that the flaw detection and rejection
limits determined for tension welds be used to form the basis
of compression weld rejection criteria. The research team
believes that it is prudent to use the same scanning procedure requirements in order to ensure that any critical flaws
be detected. The proposed acceptance criteria for tension
welds does not rely on length measurement for indications
greater than –10 dB. Therefore, purely modifying the maximum length of flaws similar to what is currently in Annex K
does not seem reasonable as this would be a very low amplitude
for automatic rejection of flaws in compression welds. Rather,
any modification to the acceptance criteria rejection limits
for compression welds should involve shifting the amplitude
limit. One option would be to shift the amplitude limit +6 dB
based on the shift that is in the existing Clause 6 amplitude
tables. However, this would be contrary to Annex K and CSA
W59:2018, which use the same amplitude limits for both compression and tension welds in cyclically loaded structures.
Table 44 displays the raster scanned results from the raster
scanned verification testing as shown in Table 36 but with
flaw rejection amplitude limits of –4 dB, –2 dB, and 0 dB.
The conventional UT rejection rate from the round robin
testing is also shown using the AWS D1.5 Clause 6 compression and tension criteria. Based on this data, setting the rejection limit to –4 dB would result in rejection of all flaws which
were rejectable by at least one technician per the round robin
testing using the conventional UT compression tables. Setting the rejection limit to –2 dB would result in acceptance
of a LOF flaw that was 0.03″ × 0.06″ and was rejectable by
60% of the technicians per the round robin testing using the
conventional UT compression tables. Setting the rejection
limit to 0 dB would result in acceptance of an HAZ crack,
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
64
Table 44. Verification testing of flaw rejection amplitude limit (compression welds).
Drawing Details
Conv. UT
Rejection Rate
Raster Scan Results
Rejection Limit
Flaw
Type
Flaw
Height
(in)
Flaw
Length
(in)
Ten
Comp
Max
Amp
(dB)
HAZ
Crack
0.18
0.52
NA
NA
8.0
Y
Y
Y
Y
Y
Y
Porosity
0.10
0.73
NA
NA
-6.2
Y
Y
N
N
N
N
-10 dB
-8 dB
-6 dB
-4 dB
-2 dB
0 dB
LOF
0.11
0.63
NA
NA
10.8
Y
Y
Y
Y
Y
Y
HAZ
Crack
0.14
0.57
NA
NA
-0.6
Y
Y
Y
Y
Y
N
Slag
0.10
0.74
NA
NA
2.3
Y
Y
Y
Y
Y
Y
Crack
0.14
0.37
NA
NA
-12.1
N
N
N
N
N
N
LOF
0.18
0.50
NA
NA
11.8
Y
Y
Y
Y
Y
Y
LOF
0.12
0.64
NA
NA
15.5
Y
Y
Y
Y
Y
Y
0.19
0.49
NA
NA
2.9
Y
Y
Y
Y
Y
Y
0.11
0.61
NA
NA
-1.0
Y
Y
Y
Y
Y
N
Toe
Crack
Toe
Crack
Slag
0.09
0.92
NA
NA
-4.2
Y
Y
Y
N
N
N
HAZ
Crack
0.14
0.45
NA
NA
3.7
Y
Y
Y
Y
Y
Y
IP
0.10
0.88
NA
NA
5.5
Y
Y
Y
Y
Y
Y
Slag
0.32
0.37
100%
60%
1.7
Y
Y
Y
Y
Y
Y
Slag
0.16
0.18
100%
100%
2.9
Y
Y
Y
Y
Y
Y
Slag
0.10
0.90
100%
100%
-1.8
Y
Y
Y
Y
Y
N
Porosity
0.09
3.31
80%
0%
-8.0
Y
Y
N
N
N
N
Toe
Crack
0.02
0.04
40%
0%
-8.4
Y
N
N
N
N
N
LOF
0.03
0.06
80%
60%
-3.6
Y
Y
Y
Y
N
N
Slag
0.06
0.03
0%
0%
-8.2
Y
N
N
N
N
N
Slag
0.17
3.61
100%
100%
0.9
Y
Y
Y
Y
Y
Y
toe crack, and slag inclusion that were rejectable by all of
the technicians per the round robin testing using the conventional UT compression tables. This is in addition to the
acceptable flaws from the –2 dB rejection limit.
The amplitude limits from the existing Clause 6 conventional UT acceptance criteria for compression welds vary
depending on the incidence angle and plate thickness. Based
on these amplitudes along with the results from the round
robin, it seems reasonable to set the ARL for compression
welds to 0 dB (i.e., equal to the SSL). While this would result
in a shift of 10 dB from the tension criteria, it seems reasonable from a workmanship standpoint. It is also recommended
that compression welds be evaluated based on the results
from the encoded line scan rather than requiring follow-up
manual raster scan of each indication. The criticality of flaws
in compression welds is much lower than in tension welds
and does not warrant the additional effort to maximize the
signal. With these recommendations, compression welds will
be essentially tested similar to how they are in the 2015 edition of Annex K since evaluation would be performed on
encoded line scans only. With this modification to the ARL
for compression welds, the new version of Annex K would
still be 5 dB more conservative than the old criteria, which
sets the ARL at +5 dB.
Based on conventional UT tables, it seems reasonable for
the EVL to be set ∼–4 dB. This is close to the old Class C limit
of –6 dB for flaws that are acceptable when 2″ or less. Therefore, it is recommended that the EVL be set at –6 dB with a
length limit of 2″. This essentially duplicates the Class C
criteria of Annex K, but are now labeled as Class B in the proposed version. Since follow-up raster scanning is not required
for compression welds, the detection limit [i.e., disregard level
(DRL)] would be set to the same amplitude as the EVL of
–6 dB. Finally, the length measurements for compression
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
65 welds can be determined using the 6 dB drop method on
the encoded line scan results similar to the current Annex K
requirements. A marked-up version of Annex K and associated commentary with the recommended changes is given in
Appendix G.
3.10 Comparison to Radiographic
Testing
Currently, most tension bridge welds are only inspected
with RT, except for fracture critical welds which are inspected
with RT and UT. Compression welds may be tested with either
RT or UT. There are select states that have replaced RT with
UT for tension welds, but this is very unique. Bridge owners
have traditionally preferred RT to UT due to the simple
interpretation of an RT image, which can be saved and easily
retrieved for permanent record. Conventional UT reports, on
the other hand, are tabulated results of the indications which
were detected by the UT technician.
RT is known to be more sensitive to volumetric flaws such
as slag inclusions and porosity while UT is more sensitive
to planar flaws such as lack of fusion and cracks. This was
apparent during this research project as most of the lack
of fusion flaws in the round robin plates were not discernable with RT while they were rejectable according to most
conventional UT and PAUT inspections.
Slag and porosity, while easily detectable with RT, had
low rejection rates with PAUT according to Annex K. Since
the rejection rates for these flaws were much greater for
conventional UT, it is likely that the poor rejection rate for
PAUT was due to differences between the acceptance criteria
of PAUT and conventional UT as well as the lack of raster
scanning with PAUT in order to maximize the amplitude.
Therefore, while slag and porosity often have low amplitude with UT (conventional UT or PAUT) compared with
planar flaws, it is not to say that volumetric flaws cannot
be detected with UT. This is apparent in the previously
reported verification testing results for flaw detection and
rejection shown in Table 42 and Table 36, where setting reasonable amplitude limits resulted in detection and rejection
of the volumetric flaws.
Digital RT images of the round robin test plates are
included in Appendix D. In addition to the intended flaws
in the weld specimens, RT images are available for two of
the three unintended flaws that were detected and rejected
if the amplitude limit was set at –13 dB. One of these indications (maximum amplitude of –10.8 dB during raster scanning) was apparent on the RT image, shown in Figure 32. It
had a maximum length of 0.03″ which would be acceptable
for all thicknesses according to the RT acceptance criteria
in AWS D1.5 Clause 6. The other unintended flaw (maximum amplitude of –11.2 dB during raster scanning) was
not apparent on the RT image since it was on the edge of
the plate.
If confronted with choosing either RT or UT, it seems that
UT would be the preferred inspection method based on the
increased sensitivity of planar flaws—which are more critical
according to FFS—along with the ability to detect and reject
volumetric flaws. It is recommended that future research
include a round robin testing program where the rejection
rate using PAUT with the proposed revisions to AWS Annex K
would be compared to the rejection rate for RT. Performance
qualification of PAUT technicians requiring the detection
and rejection of critical planar and volumetric weld flaws
during an independently administered practical examination would provide additional verification that PAUT is providing adequate sensitivity to critical flaws of all types.
3.11 Technician Performance
Qualification
3.11.1 Current AWS Requirements
The round robin testing showed that there was a large
amount of variability in inspection results using the PAUT
Annex K and conventional UT codes. While some of this variability may be due to differences in inspection equipment,
equipment settings, and scanning procedures, a large portion
of this variability is due to human factors and inconsistencies.
For instance, it is likely that the large amount of variability in
the reported location of the same flaw was primarily caused
by poor calibration of encoders. It was noted that technicians
Unintended Flaw
Figure 32. RT image of unintended weld flaw.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
66
would often consistently report multiple flaws either to the
left or right of their actual location. This offset in flaw location was sometimes quite large, resulting in a large number
of detected flaws which did not meet the API RP 2X requirements for reported flaw location.
The inspection variability due to human factors would likely
be improved if PAUT technicians were required to pass practical examinations that were administered by independent entities. These proposed examinations would include inspection
of flawed weld specimens using AWS D1.5 Annex K. The test
plates should include known flaws that are comparable to the
critical flaw size used to determine the acceptance criteria.
Since AWS D1.5 and ASNT SNT-TC-1A do not currently
have any guidance on critical weld flaw size, it is unknown
what size flaws are included in the practical examinations
performed in accordance with ASNT SNT-TC-1A [48]. In
addition, ASNT SNT-TC-1A only requires that a minimum
of one flawed specimen be used for the practical examination
without any guidance on the number of flaws, type of flaw,
orientation of flaw, or requirement for blank specimens. It
states that the flawed specimen should be representative of
the component that would be tested, but that interpretation
is left to the ANST Level III. Finally, no minimum requirements for passing the practical examination are provided.
It states that at least ten different checkpoints requiring an
understanding of NDT variables and the employer’s procedural requirements should be included in the practical exam.
It also states that the candidate should detect all discontinuities and conditions specified by the NDT Level III. Finally, it
notes that while it is normal to score the practical on a percentile basis, practical examinations should contain checkpoints or gateway tasks that failure to successfully complete
would result in failure of the examination.
The 2016 edition of SNT-TC-1A included a sample checklist
for guidance on the development of practical examinations.
This checklist is not specific to any method or level and may
be modified as needed in accordance with the Level III. The
sample checklist includes a possible breakdown of a scoring rubric and sample limits on flaw detection, false calls,
and flaw evaluation. The ten categories listed in this sample
checklist include
1.
2.
3.
4.
5.
6.
7.
8.
9
10.
Knowledge of NDT Procedure
Equipment and Material
Test Specimen Care and Custody
Operations
Detection of Indications
Interpretation of Indications
Evaluation of Indications
Documentation and Records
General Health and Safety
General Observable Conduct
Since there is no guidance on the flaw size in the practical exam specimen, the weld flaws included in the specimen
could be extremely large such that anyone who understands
the very basics of PAUT inspection would be able to detect
them and would pass the practical exam. In this case, the
examination is not performing the intended function of
testing the competency of the individual. This would be like
giving a structural engineer a single question exam on steel
design where they only had to compute the tension stress by
dividing the given force by the given area and deeming that
they are now qualified to design a bridge.
ASNT SNT-TC-1A states that the technician should demonstrate familiarity with and ability to operate the NDT equipment, record, and analyze the resultant information to the degree
required. The “degree required” statement is vague and leaves the
decision on whether the technician has adequate performance
completely to the ASNT Level III administering the exam. It also
states that the Level II PAUT technician should detect all discontinuities and conditions specified by the NDT Level III. There are
no requirements on the accuracy of flaw location measurements
or limitations on the number of false calls.
In the opinion of the authors of this report, discussion is
merited on the self-policing of the NDT industry according
to ASNT SNT-TC-1A. There are no specific requirements on
the difficulty of the practical test or on the method for
grading the practical test. It is the NDT firm’s advantage to
have as many technicians pass the exam and be available for
inspection duties. It is also difficult for NDT firms to have
a large number of samples available or to have specimens
which have not been used for previous tests. The test specimens
may be reused from technician to technician within an NDT
inspection firm or even reused for reexamination of candidates who previously failed (or passed) the exam.
ASNT SNT-TC-1A includes no discussion on characterization of flaw type, yet AWS D1.5 Annex K requires that flaws
characterized as cracks be rejected. In other words, the current training and certification program for PAUT technicians
does not include any requirements on the ability to characterize flaws, but it is expected that these technicians will be able
to accurately characterize the flaw type when they perform
weld inspections per AWS D1.5 Annex K. (As noted, the round
robin phase of the research showed the current workforce has
limited reliability in this skill.)
The AWS D1.1 proposal for PAUT inspection has modifications to the personnel qualification requirements, including doubling the minimum number of hours of work time
experience in PAUT from 160 hours to 320 hours and requiring that the practical exam consist of at least two flawed
specimens representing joint types to be examined with
each specimen containing a minimum of two flaws. The
research team believes that doubling the minimum number
of hours of work time experience is unlikely to result in a
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
67 large improvement in PAUT inspection quality. In fact, it
may actually have the opposite effect as a technician who is
not properly performing any given task will become more
entrenched in the wrong practice and become more confident
that he or she is actually doing it correctly. In short, requiring
a PAUT technician (or any individual performing any task) to
perform additional work time experience does not necessarily
mean that the technician will perform “better,” as it is unclear
whether the technician is performing the inspection correctly
in the first place. For example, a technician who has not properly demonstrated that they can detect and reject critical weld
flaws but has 320 hours of work experience incorrectly performing PAUT inspections is unlikely to be any better than he
was after he completed the first 160 hours of incorrect PAUT
inspections. Rather than doubling the required work time
experience, improving the meaningfulness of the practical
examination and setting more defined requirements for passing the practical examination would likely result in greater
improvement in technician performance.
While there is merit to increasing the number of flawed
specimens and the number of flaws tested during the practical examination, it does not seem reasonable to require a
minimum number of flaws per flawed specimen as this only
provides the candidate with a minimum number of hits per
plate that they need to find. It would be much better to have a
random set of specimens with various number of weld flaws
mixed with blank specimens which do not have any flaws.
This way the candidate does not know how many flaws there
are per plate and does not expect that there should be a flaw
in each plate. This results in a much more realistic practical
exam since the number of flaws in a weld is always unknown
in an inspection. Other guidance on the development of a
performance test to evaluate reliability of a nondestructive
testing system is given in MIL-HDBK-1823A [49].
Some states have started to recognize the need for improved
practical examination of UT technicians. For example, New
York State Department of Transportation (NYSDOT) has an
Ultrasonic Testing Technician Program included in their Steel
Construction Manual [50]. This program requires prospective UT technicians who wish to be certified by NYSDOT to
pass examinations provided by NYSDOT. These examinations
include an open and closed book exam worth 25% each as well
as a practical exam worth 50%. NYSDOT also keeps a list of
Certified UT Technicians available on their web page. One of
the administrators of this UT certification program stated that
a large percentage of prospective UT technicians have failed
their state’s exam. In fact, they are lessening the required time
between initial test and the retest in order to make it easier for
retesting. This case study highlights the need for independent
examination of UT technicians.
Texas Department of Transportation (TxDOT) [51] also
includes a hands-on examination administered in-house in
addition to the requirements of the AWS code and employer’s
Written Practice. A TxDOT official who administers this
exam was contacted to confirm the type and size of flaws used
in this exam. TxDOT has a specific plate for each geometry,
including T and corner joints and a thickness transition butt
weld. Each plate has multiple flaws that are of various types,
including both planar and volumetric flaws. These flaws are
typically Class A rejectable defects in accordance with Clause
6 tension weld conventional UT tables.
While the NYSDOT and TxDOT programs are for additional practical examination of conventional UT technicians,
Florida Department of Transportation (FDOT) is currently
implementing a program to test PAUT technicians. This program involves practical examination of each PAUT technician using an in-house test block and QA inspector trained
in PAUT. The flaws chosen for this block are specific to the
response of FDOT’s internal research. An FDOT official
stated that this qualification was deemed necessary based on
the level of training that technicians were receiving in typical
PAUT training courses. Since FDOT is interested in replacing
RT with PAUT, the selection of flaws for the practical examination was determined to be critical and was one of the reasons that this qualification is performed in-house.
3.11.2 AWS Recommendations
As PAUT use becomes more prevalent, it will be important
for the PAUT technician qualification requirements to be standardized across individual agencies. In speaking with these
states, it became clear that having an independent central organization that would administer a reasonable practical examination using Annex K would be beneficial. By standardizing
the practical examination, PAUT technicians would not need
to take a separate exam to satisfy the individual requirements
of each agency. Rather than having many separate examinations with one flawed plate, a standardized examination could
be much more thorough and still take less time overall.
A good example of a thorough, standardized qualification
requirement is the Certified Welding Inspector (CWI) examination. Instead of having separate exams for each agency, prospective CWIs only need to qualify through AWS. The AWS
CWI examination process includes a three-part examination
that extends over a full day. Two parts of the exam cover welding background (closed book) and code requirements (open
book) while the third part is a practical examination using weld
specimens. The CWI examination is set by AWS and proctored
at defined locations throughout the year. The CWI examination
has a controlled level of difficulty and clearly defined expectations that covers visual examination of welds for all agencies.
The authors recognize that requiring independent practical examination of PAUT technicians will result in additional
cost but believe that this may be necessary to implement the
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
68
removal of RT requirements in lieu of in-depth PAUT inspection. The researchers are confident that a meaningful practical examination could be performed in a single day.
The research team believes that additional research should
be performed to develop a PAUT technician qualification
program which utilizes independent and consistent evaluation of technicians. This study could compare practical
examination results to shop performance in order to develop
a quantitative practical examination that will adequately
evaluate candidates for bridge weld inspections. Based on
the research and what was discussed previously, the performance qualification requirements in the following sections
are proposed.
3.11.2.1 Practical Examination Parameters
The practical examination should involve meaningful performance testing that is conducted by an independent party.
The research team recommends that, optimally, the test
would be administered by a third party that has developed a
realistic test procedure that is acceptable to all agencies.
The practical examination should evaluate detection, location, and rejection of realistic weld specimens with critical
flaws. In addition, the practical examination should include
evaluation of following parameters:
• Familiarity and application of Annex K requirements
• Development of adequate scan plan and documentation
•
•
•
•
of essential variables
Use of proper equipment
Proper calibration for reference sensitivity and acoustic
properties
Proper application of the two-part inspection procedure
(line scan and follow-up raster scanning)
Proper documentation and reporting
3.11.2.2 Specimen Details
The specimens used in the practical examination should be
fabricated from steel that have acoustic properties representative of the typical steels encountered in bridge weld inspection but also include some specimens in which calibration
adjustments are required to account for variation in acoustic
properties. The specimens should be representative of the configuration of the welds that will be inspected by the technicians
during future applications of Annex K. This includes thickness
transitions and weld reinforcement if both will be encountered.
The number of flaw specimens shall be large enough such
that the number of flaws per plate will be varied, and blank
specimens should be included in the lineup of plates to be
tested. The technicians should be instructed to detect and
report all relevant indications and should be informed that a
plate may not necessarily contain any flaws. The research team
believes that, based on the previous research, including the
round robin, at least four weld specimens should be used with
a weld length of approximately 18 inches. Weld specimens that
are too short may not adequately capture errors from encoded
line scanning. More weld specimens may be necessary in order
to adequately include possible weld configurations and acoustic properties (proper calibration practices) that technicians
may encounter during weld inspection.
The number of flaws included in the weld specimens should
be adequate to verify satisfactory performance. Both volumetric and planar discontinuities should be included in the set of
weld flaw specimens. It is recommended that at least one slag
inclusion, porosity grouping, lack-of-fusion flaw, and vertical planar flaw (e.g., crack or incomplete penetration) should
be included. Variations in the flaw size and through-thickness
location will also be necessary in order to verify that the scanning procedures for flaw detection and rejection are properly being followed. The absolute minimum number of flaws
included in the practical exam is recommended to be five. Use
of more flaws and weld specimens would allow for variation
in flaw size and location (through-thickness depth, tilt, and
location along weld axis), which would provide additional
information on the scatter of results from human factors.
3.11.2.3 Pass/Fail Criteria
It is believed that the formulas and minimum performance
levels included in API RP 2X are appropriate for the bridge
industry for evaluating flaw detection, sizing, and location
during the practical examination. API RP 2X includes minimum recommended performance levels for UT technicians
using the following two formulas:
P=
Lc
× 100
La
Lc
Lf
R =    1 −  × 100
 L1  
L1 
Formula 1
Formula 2
Where:
P =percentage of actual reflectors correctly detected and
sized, 0 to 100
R = overall rating including penalty for false calls, 0 to 100
La = length of actual reflector contained in the test plate
Lc =credited length for indications that have been correctly sized and located (credit is given for the lesser
of the reported length or actual length of the flaw.)
L1 =accumulative length of all indications by the technician, right or wrong
Lf =accumulative length of indications above the stated
disregard level where no reflector exists
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
69 To be correctly sized, this document recommends that
the reported dimensions be within a factor of two of true
dimensions (one-half to twice the actual dimension). To
be correctly located, this document recommends that the
centerline of the reported indication be within the boundary
of the actual indication or within ½ inch of the actual centerline of the indication (whichever is greater). API RP 2X suggests that a score of 70 or above for Formula 1 and a score of
50 or above for Formula 2 be used as minimum performance
for ultrasonic technicians.
When the round robin test results were used to calculate the performance of each conventional UT and PAUT
technician, two of the five conventional UT technicians and
zero of the four PAUT technicians met these levels. The
primary reason for the PAUT technicians not meeting the
minimum performance was due to errors in flaw location.
As stated, this is thought to be due to issues with the use of
encoders. Since the flaw has to be detected in order to count
towards the performance requirements, the lower sensitivity with Annex K prior to the proposed modifications (no
separate flaw detection limit for the line scans) will have
also contributed to some of the low scores. The primary
reason for the conventional UT technicians not meeting
the minimum performance was due to poor accuracy for the
flaw length measurement and some misses of long scattered
porosity. The issues with flaw length measurement are likely
due to the manual method for length measurement with
conventional UT, which lends itself to overestimating the
actual length.
While the minimum scores provided by API RP 2X could
be used to evaluate flaw detection and flaw location, the
pass/fail criteria of the practical examination should also
evaluate the following items in determining an overall
practical examination score:
• Familiarity and application of Annex K requirements
• Development of adequate scan plan and documentation
of essential variables
• Use of proper equipment
• Proper calibration for reference sensitivity and acoustic
properties
• Proper application of the two-part inspection procedure
(i.e., line scan and follow-up raster scanning)
• Evaluation of indications using proper acceptance criteria
• Proper documentation and reporting
Using a checklist similar to that provided in the 2016 edition
of ASNT SNT-TC-1A, a rubric should be developed that incorporates each of these items. By assigning points to each item,
along with flaw detection and location criteria, technicians
passing the practical examination will demonstrate adequate
comprehension and application of the Annex K inspection
procedure. Suggested performance testing guidance has been
added to the proposed modifications to Annex K found in
Appendix G of this report.
3.12 Flaw Sizing Acceptance Criteria
for Alternative UT Methods
An acceptance criteria based on flaw height and length sizing
is provided as an alternative method to the amplitude approach
provided previously. The alternative method requires that the
PAUT technician develop a written procedure according
to specified requirements as well as demonstration of the
procedure through performance testing on samples of similar material and with flaws similar to the rejectable size.
This allows for other advanced ultrasonic methods such as
FMC/TFM PAUT or TOFD to be used to inspect bridge welds
provided that they can detect and reject critical weld flaws.
It is envisioned that the requirements for alternative ultrasonic systems and the accompanying acceptance criteria
based on flaw sizing would be included in the main body
of AWS D1.5, rather than included in Annex K. Annex K
includes all necessary requirements and acceptance criteria
for application of encoded line scanned PAUT. Annex K does
not rely on accurate flaw height sizing and, therefore, does
not require development and qualification of an individual
flaw sizing procedure for the specific equipment and application method.
CSA W59-18 [11] includes minimum requirements
for alternative ultrasonic systems in the main body of the
code. Alternative ultrasonic systems include encoded PAUT
and TOFD in W59, as only manual raster scanned PAUT
are allowed to replace conventional UT without additional
performance testing. Because AWS D1.5 Annex K covers
encoded PAUT using amplitude and length acceptance criteria, it is recommended that the additional requirements
for alternative UT methods in AWS D1.5 apply to any methods
that do not fit within the requirements of conventional UT
Clause 6 or Annex K.
3.12.1 Minimum Requirements
The following requirements from CSA W59 are recommended for incorporation into AWS D1.5:
• Written agreement by the engineer and contractor prior
to the examination allowing the use of the alternative
inspection method
• Certification of Level II or III in accordance with ASNT
SNT-TC-1A for the specific method, if applicable; for
instance, ASNT SNT-TC-1A includes qualification requirements for TOFD, but does not include additional qualification requirements for FMC/TFM PAUT
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
70
• Documentation of inspection procedures in writing in
accordance with recognized standards and accepted in
writing by the engineer
• Written procedures that contain at minimum the following information:
–– Specific operator training requirements
–– Types of weld joint configurations to be examined
–– Acceptance criteria
–– Type of UT equipment (manufacturer and model
number)
–– Type of transducer, including frequency, size, shape,
angle, and type of wedge
–– Scanning surface preparation and couplant requirements
–– Type of calibration test block(s) with appropriate reference reflectors
–– Method of calibration and calibration interval
–– Method for examining for laminations prior to weld
evaluation
–– Scanning pattern and sensitivity requirements
–– Methods for determining discontinuity location, height,
length, and amplitude level
–– Transfer correction methods for surface roughness,
surface coatings, and part curvature, if applicable
–– Method of verifying the accuracy of the completed
examination; this verification may be by reexamination
using UT by others, other NDE methods, macroetch
specimen, gouging, or other visual techniques accepted
by the engineer
–– Documentation requirements for examinations, including any verification performed
–– Documentation retention requirements
–– Demonstration of the system in order to achieve engineer’s approval
� System demonstration on one or more demonstration blocks simulating the weld
� Demonstration of adequate and repeatable detection of typical weld flaws and useful to set threshold
parameters
–– Acceptable performance of system demonstration should
include:
� Detection of all of the flaws in the demonstration
block(s)
� Recorded flaw sizes for critical weld flaws with reported
flaw size that exceeds the acceptance criteria limits
� Recorded flaw sizes for subcritical weld flaws with
reported flaw size that is reasonably accurate
3.12.2 Acceptance Criteria
The acceptance criteria used for the alternative UT methods, which uses measurements of flaw height and length,
should be comparable to the acceptance criteria used in the
recommendations for Annex K acceptance criteria for planar
Table 45. Alternative UT methods acceptance
criteria for planar embedded flaws.
Flaw Height
(in)
0.06
0.07
Flaw Length
(in)
1.00
0.28
0.10
0.16
0.14
0.14
Interpolated Length (in) for
Intermediate Height (in)
For 0.07<H<0.10,
L=0.56-4.0*H
For 0.10<H<0.14,
L=0.21-0.5*H
and volumetric flaws. As described in Section 3.7, the acceptance criteria for Annex K was developed using the critical
flaw sizes computed using fracture mechanics.
As shown in Table 33, the –10 dB amplitude limit for flaw
rejection was based on CIVA analysis of a 0.05″ × 0.10″ planar
surface flaw. This flaw was comparable to the 0.06″ × 0.08″
critical planar surface flaw for a 4 ksi stress range thickness
transition weld and an 8 ksi stress range equal thickness weld.
The controlling planar embedded flaw had an amplitude of
–8 dB, according to CIVA, using a 0.15″ × 0.15″ planar embedded flaw. This flaw was comparable to the 0.17″ × 0.17″ critical
planar embedded flaw for a 4 ksi stress range thickness transition weld and the 0.14″ × 0.14″ critical planar embedded flaw
for an 8 ksi stress range equal thickness weld.
For the acceptance criteria used for the alternative UT
methods, it is recommended that the flaw height and length
limits be provided for various a/c ratios. The results for the
critical planar flaws noted above were summarized into
Tables 45 and 46. Table 45 includes a recommended acceptance criteria for planar embedded flaws measured using
alternative UT methods. Table 46 includes a recommended
acceptance criteria for planar surface flaws.
As shown in Table 34, the –13 dB amplitude limit for rejection of flaws 1″ long or greater was based on CIVA analysis
of a 0.125″ diameter embedded volumetric flaw and a 0.08″
diameter surface volumetric flaw. The embedded flaw was
comparable to the 0.14″ diameter critical volumetric embedded flaw for an 8 ksi stress range equal thickness weld. The
surface volumetric flaw was comparable to the 0.08″ × 0.11″
critical volumetric surface flaw for a 4 ksi stress range in a
thickness transition weld and a 0.06″ × 0.08″ critical voluTable 46. Alternative UT methods acceptance
criteria for planar surface flaws.
Flaw Height
(in)
0.02
0.03
0.06
Flaw Length
(in)
1.00
0.20
0.09
Copyright National Academy of Sciences. All rights reserved.
Interpolated Length (in) for
Intermediate Height (in)
For 0.03<H<0.06,
L=0.31-3.67*H
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
71 metric surface flaw for an 8 ksi stress range in an equal
thickness weld.
Since these flaws are the same sizes as the critical planar flaws, it is recommended that the limits provided in
Table 45 and Table 46 also be used for the maximum size
of an individual volumetric flaw (i.e., maximum slag inclusion or pore). In addition, a maximum length of 1″ is recommended for scattered indications such as a group of porosity.
This is similar to the length requirement recommended for
Annex K for flaws greater than –13 dB but less than –10 dB.
Since the inputs to the FFS study were for tension welds. The
limits given in Table 45 and Table 46 will be overly conservative
for compression welds. The recommendation for the Annex K
amplitude limits for the compression acceptance criteria uses
a 0 dB amplitude. This amplitude approximately correlated
with the maximum amplitude of vertical 0.10″ × 0.10″ planar
surface flaw and a vertical 0.20″ × 0.20″ planar embedded flaw
from CIVA analysis. Note that this amplitude was based on traditional conventional UT amplitude limits rather than fracture
mechanics. For compression welds, a 50% increase in the critical flaw height and length given in Table 45 and Table 46 seems
reasonable to compare with the Annex K acceptance criteria.
Another strategy would be to set this limit based on a strength
requirement and an acceptable amount of unfused material.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
72
CHAPTER 4
Conclusions and Suggested Research
4.1 Conclusions
NCHRP Project 14-35: Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using
Enhanced Ultrasonic Methods had the objectives of developing guidelines to evaluate CJP welds in steel bridges based on
updated acceptance criteria and to develop proposed modifications to AWS D1.5. While AWS D1.5 currently includes
PAUT inspection procedures in Annex K, these acceptance
criteria were workmanship-based and were carried over from
previous D1.1 conventional UT methods. AWS D1.5 did not
provide means for alternative methods such as FMC/TFM
PAUT or TOFD, which are suited for evaluation of flaw criticality based on measurements of flaw size rather than amplitude responses.
The research has shown that (1) the critical flaw size
of bridge welds could be developed using FFS, (2) a large
amount of variability was possible when weld inspections
were performed using current AWS D1.5 conventional UT
and PAUT scanning procedures and the current workforce,
(3) computer modeling could be used to evaluate ultrasonic
responses of weld flaws and reference reflectors, (4) acoustic
properties of bridge steels may vary widely and may not be
isotropic, (5) revised acceptance criteria for Annex K could
be developed to detect and reject critical weld flaws utilizing
amplitude-based criteria, and (6) additional technician performance requirements including independent practical
examination were necessary.
Major findings of this research were that the current scanning procedures and acceptance criteria in Annex K did not
correlate to traditional limits used in conventional UT per
Clause 6 and were not adequate for rejection of critical weld
flaws according to FFS. This research also found that differences in acoustic properties between the calibration block and
test object could result in significant error in reference sensitivity for frequencies that were allowed in Annex K. The current version of Annex K does not include any requirements on
calibration block acoustic properties. Variations in shear wave
velocity were also found to be significant for common grades
of bridge steels. These variations resulted in significant error
in beam refraction angle, which could result in inaccurate flaw
location and significant loss of amplitude. These variations
affect both conventional UT and PAUT. Proposed revisions to
Annex K were provided in this report that would account for
these differences, but similar revisions should be included in
Clause 6 for conventional UT.
The final product of NCHRP Project 14-35 was proposed
revisions to AWS D1.5 Annex K for improved flaw detection
and rejection. These revisions include minimum requirements
for technician qualification, requirements on the acoustic
properties of calibration blocks in order to represent the conditions found in the test object, requirements on the scanning
procedure and sound coverage, and requirements on acceptance criteria to detect and reject critical weld flaws.
Recommendations were also provided which would allow
alternate UT methods to be used in lieu of PAUT or conventional UT. Use of these methods relies upon written agreement by the engineer and contractor along with procedure
development and demonstration on weld flaw specimens.
Limits on acceptable flaw sizes that could be incorporated
into an acceptance criteria based on flaw size measurements
were also provided.
4.2 Suggested Research
While NCHRP Project 14-35 has resulted in important findings and recommendations for modifications for AWS D1.5
regarding application of PAUT for the inspection of bridge
welds, additional research is suggested which could aid in
application of these recommendations. Five topics of suggested research are proposed: (1) performing a round robin
testing program to compare inspection results using RT to
inspection results using PAUT in accordance with the revised
version of Annex K, including the proposed modifications;
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
73 (2) developing a performance-based qualification program
for PAUT technicians and verifying improvement in inspection results; (3) developing specific scan plan recommendations for probe selection and line scan index offset for
typical weld geometries; (4) collecting additional data on
variability of acoustic properties for steel bridge base metals;
and (5) developing automated scanning and data processing
techniques to minimize variability due to human factors.
The first research topic will compare round robin inspection results of RT and PAUT using the revised version of
Annex K. While this research project used computer modeling and experimental test results to develop and verify the
revisions to Annex K, the round robin results would be able to
collect information on human factors and scatter in results.
This study would aid in adoption of the revised Annex K
inspection procedure to be used in lieu of RT by providing a
direct comparison of inspection results using both methods.
The second research topic will involve development of
PAUT technician qualification requirements, which will adequately demonstrate use of the revised Annex K procedure.
This research would involve the establishment of the number,
size, and type of flaws needed for a meaningful practical
examination. This research will also provide verification
whether the PAUT inspection quality will meet or exceed RT
after requiring proper training and performance testing of
PAUT technicians along with the revisions to Annex K.
The third research topic will develop specific scan plan
recommendations for typical weld geometries. Currently,
there is no ANST Level III for PAUT; therefore, the ASNT
Level II PAUT technicians are responsible for developing the
scan plan. This includes probe selection, index offset location for line scans, and focal law configurations. As shown
in NCHRP Project 14-35, the probe frequency and aperture
should be properly selected in order to result in optimal
inspection results. For instance, inspections of thick plates
with long sound paths may require use of lower frequency
and larger apertures in order to account for loss of amplitude
due to attenuation and beam spread. This research has also
shown that the amplitude response of flaws will be greatly
influenced by probe location. Since typical bridge welds
utilize very similar geometries based on the plate thickness,
recommendations to probe parameters and index positions
could be tabulated for typical welds. This would cut down
on the effort for scan plan development and would result in
more consistent inspections from technician to technician
within the QA/QC process.
The fourth research topic will collect additional data on
the variability of acoustic properties for steel bridge base
metals. This research topic would involve collecting data on
the variability of acoustic properties of applicable steel bridge
grades along with the variability of acoustic properties at different locations within the same heat of steel. This research
topic will aid in determining the overall scatter of typical
steels and in developing refined calibration standards. This
project should include combinations of possible heat treatment within each grade, including A709-50CR (i.e., A1010)
since the acoustic properties of this mild stainless steel were
not investigated during NCHRP Project 14-35.
The fifth research topic will develop automated scanning
and data processing techniques to minimize variability due
to human factors. This research topic would involve evaluating automated scanners, which would result in consistent
data collection and probe manipulation. It would also evaluate techniques for automatic data processing in order to
consistently evaluate scan results according to the proposed
acceptance criteria.
Copyright National Academy of Sciences. All rights reserved.
Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
74
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
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Acceptance Criteria of Complete Joint Penetration Steel Bridge Welds Evaluated Using Enhanced Ultrasonic Methods
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Fixing America’s Surface Transportation Act (2015)
Federal Highway Administration
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